STRUCTURE magazine | November 2023

STRUCTURE NOVEMBER 2023

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INSIDE: Design of Damped Moment Frames 12

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Contents Cover Feature

12 PRESCRIPTIVE DESIGN OF DAMPED MOMENT FRAMES

By Nathan Canney, Ph. D., P. E., Alan Klembczyk, Konrad Eriksen

Fluid Viscous Dampers, or simply “dampers,” effectively dissipate seismic energy to reduce steel and foundation costs in new buildings and improve structural resilience, especially for critical facilities.

NOVEM BER 2023

Features 22 VISION NORTHLAND - MINNESOTA HOSPITAL DESIGN BRACES FOR WINTER AND THE FUTURE

By Colleen F. Blackwell, P. E., Stephen G. Bartal, P. E.

Essentia Health’s Vision Northland project, located on the shores of Lake Superior in Duluth, Minnesota, includes approximately 928,000 square feet of new construction and roughly 120,000 square feet of renovation to existing facilities and is scheduled to be completed by the end of 2023.

28 TERMINUS

By Ilana Danzig, P. E., S. E., P. Eng, Struct Eng, M. Eng

The completion of a 5-story mass timber commercial office building named Terminus in the city of Langford, British Columbia, on Canada’s west coast, marks an exciting first in the world of seismic systems for timber structures.

Columns and Departments 9 Editorial Mentoring is for the Birds

By Cervente Sudduth P. E., ENV SP,

40 Iconic Structures San Francisco–Oakland Bay Bridge

By Roumen V. Mladjov, S. E., P. E.

64 Insights HSS Availability

By Beth Suminski, P. E., S. E., and Cathleen Jacinto, P. E., S. E.

10 Infocus Improving Engagement and Career Longevity in Structural Engineering

By John A. Dal Pino

18 Structural Design Design of Stainless Steel Bolted Connections in Accordance With the Recently Published ANSI/AISC 370

By Francisco Meza, Ph.D. and Nancy Baddoo

36 Code Updates 2024 IBC Significant Structural Changes

By John “Buddy” Showalter, P. E. and Sandra Hyde, P. E.

46 Historical Structures 19th Century Mississippi River Bridges #9

By Frank Griggs, Jr., Dist. M. ASCE, D. Eng,. P. E., P. L. S.

48 Structural Connections Vibration of Steel Joists with Flush Frame End Connections

By David Samuelson, P. E., Brad Davis, Ph. D., S. E., and Thomas M. Murray, Ph. D., P. E.

In Every Issue 3 Advertiser Index 55 Software Updates Guide 58 NCSEA News 60 SEI Update 62 CASE in Point

On the Cover: Voluntary seismic retrofit of a Pre-Northridge Steel Moment Frame building in California using fluid viscous dampers.

Photo Courtesy of Alejandro Velarde, (www.alejandrovelarde.com).

Publication of any article, image, or advertisement in STRUCTURE® magazine does not constitute endorsement by NCSEA, CASE, SEI, the Publisher, or the Editorial Board. Authors, contributors, and advertisers retain sole responsibility for the content of their submissions.

STRUCTURE magazine is not a peer-reviewed publication. Readers are encouraged to do their due diligence through personal research on topics.

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EDITORIAL Mentoring is for the Birds

By Cervente Sudduth P. E., ENV SP,

E

arly in my college career I recall taking several required arts and science courses and certain required elective courses. A specific memory comes to mind regarding a book report the class was assigned midway through the semester in which I had to write an expository essay. I must have done an awesome job on the essay as I received an “A.” The course TA called me to her office to discuss the essay. She told me it was well written and that she was very impressed. She then asked what my course of study was at the university. After I said Engineering, she suggested I think about changing my major to English because she felt I had a really great future in the field given the essay I had written. As a young student and throughout my college career, this comment always resonated with me. I was never able to grasp why or how a teacher would coach/mentor a student away from set goals.

No matter how sound her advice was, it did not waiver my journey to become a civil engineer.

Why is this story important? Because Mentors provide a variety of functions that support, guide, and counsel those that are less experienced. The influence we exert as seasoned individuals can easily sway or misguide students as well as emerging engineering professionals.

According to the Psychologist Daniel Levinson in his book The Season’s of a Man’s Life, a young adult early in their career is likely forming an occupational identity along with a dream, professional relationships, and mentor relationships.

It is at this point a young person or professional is establishing “Role Identity vs Role Confusion” and “Professional Intimacy vs Isolation.” During this period of professional development, mentorship is imperative.

How do we define mentorship? Here are two definitions: • Mentoring is a reciprocal and collaborative at-will relationship that most often occurs between a senior and junior employee for the purpose of the mentee’s growth, learning, and career development.

• Mentorship is a professional, working alliance in which individuals work together over time to support the personal and professional growth, development, and success of the relational partners through the provision of career and psychosocial support.

Mentoring should be the cornerstone of how we operate as structural engineers. As I prepared to write this editorial, I found it insightful to ask our STRUCTURE magazine

staff what mentoring and mentorship means to them. I was not at Through enabling others, mentors satisfy all surprised by the an important generative need. Having the responses. opportunity to positively impact the trajectory • Mentoring bridges of someone else’s career path by providing the gap between sound effective counseling and coaching has an what is learned in school and unparalleled feeling of self-gratification. what is practical application in the industry.

• Providing guidance to new staff to assist in understanding mentor/mentee relationship, the benefit predomcompany policies, procedures, and ways of inantly favors the mentee. Boston University’s executing work tasks.

Kathy Kram suggests the contrary: Through • Having a resource available to you that can enabling others, mentors satisfy an important genprovide knowledge and expertise in an area erative need. Having the opportunity to positively specific to an assigned task. impact the trajectory of someone else’s career path by Additionally, I asked the staff, on a scale of 1 to 10 providing sound effective counseling and coaching how important mentorship was in the workplace has an unparalleled feeling of self-gratification. As (1 is absolutely not important; 10 is absolutely structural engineers, it also is our duty. important). It was 10s across the board. The staff Mid-way through my career, I was a mentor overwhelmingly confirmed having a mentor or and coach to my structural team. I had a young culture of mentorship was vital to workplace hap- engineer leading all design and drawing production piness and success. on a project for the first time. He had experience in Randstad, a global HR service firm, found that all facets of building design, but he had not been employees participating in a mentor program were provided with a chance to lead. The day our team 49% less likely to leave their organization. This submitted 100% bid documents, he came to my equated to a cost savings of $3,000 per employee office and said, “I want to thank you for giving per year, according to Matthew Reeves, CEO of me wings.” He then gave me a card expressing his Together. In the same study, a Gallup poll found gratitude. Twenty-five years later, I still have the 68% of employees in the U.S. were disengaged card displayed in my office on my mantle. While with their work. Reeves goes onto say, Mentorship my colleague achieved a milestone in his career programs and an organizational culture that that day, I received a sense of accomplishment supports and encourages mentorship enhances right alongside him. employee engagement because: Why is Mentoring For the Birds? Mentoring is • It provides more opportunities for training and for the Birds because no one knows to what heights development by tapping into the knowledge they can soar until they spread their wings and fly. of your more senior employees.

This quote is engraved in a photo of a bald eagle • Mentorship gives employees a voice to speak flying high in the sky that I keep at the edge of with leadership, thus breaking down barriers my desk.■ to communication.

• Both mentor and mentees are given the opporCervente Sudduth P. E., ENV SP, is the President tunity to prove themselves by putting into of DuBois Consultants, Inc. located in Kansas City, practice what they discuss during sessions.

MO. He is an active member on the NCSEA SE3 • Engagement is closely tied with working relaCommittee (Committee Liaison), serves as a Director tionships. Mentorship builds social ties. on the Board of Directors for NCSEA, and has a M. • Mentorship holds mentors and mentees S. in Civil Engineering from the University of Missouriaccountable for commitments they make to Kansas City and a B. S. in Civil Engineering from the one another.

University of Missouri-Columbia.

There’s a common misconception that in a N OVE M B ER 2023

9

INFOCUS Improving Engagement and Career Longevity in Structural Engineering Let’s all agree to identify and remove obstacles.

By John A. Dal Pino

O

ver the past few months, I have written about employment in the structural engineering profession. The first article in the February 2023 issue dealt with the apparent shortage of structural engineers, including an analysis of the potential causes. I concluded that engineers are just as responsible for the situation as anyone or anything else. The second article in the July issue dealt with leadership and mentorship.

Again, I stated that we can be our worst enemies but that there is an opportunity for improvement.

I was motivated to continue thinking on this theme when I saw an announcement on LinkedIn celebrating June 23 as International Women in Engineering Day. Since STRUCTURE is a monthly magazine with a long lead time in article preparation, that day has passed, but let’s celebrate again anyway. Who doesn’t enjoy tipping back a flute or two of champagne with co-workers?

When I started my career in structural engineering, there were far fewer women in the profession than now. We have come a long way, but I don’t know if this change has been unique to structural engineering. I don’t have any statistics to back me up, but I suspect that there were far fewer women in traditionally male-dominated professional careers 50 years ago. The societal changes in the late 1960s and 1970s were responsible for pushing forward and expanding higher educational opportunities. It has taken several decades for the change to occur and for us to really “see” the impact. Some fields have seen such significant moves toward equality that young adults, in particular, may forget this has not always been the situation. Perhaps the more male-dominated fields of STEM and construction are taking longer since there was and is more that needs to change.

If you want a real shock to your system, read a copy of a magazine from the 1920s. I recently purchased a copy of The Atlantic magazine from October 1929, published just before the Great Depression started. Every advertisement was targeted at men. Every description used male pronouns exclusively. It was as if “she” and “her” hadn’t been invented yet. Despite my love of fine champagne, I am looking forward to the day when celebrating women in engineering at the office happy hour doesn’t seem to be necessary any longer, non-binary pronouns are used interchangeably without notice, and we have declared victory and moved on to the next worthy and pressing cause.

As I mentioned in the July 2023 article, several surveys by the SE3 Committees of NCSEA and the Structural Engineers Association of Northern California (SEAONC) generated data from working professional engineers on career satisfaction, career development,

10 STRUCTURE magazine

pay and benefits, work-life balance, etc. Amongst the many findings, I want to focus now on the following:

• 55% of all respondents (65% of women) had considered leaving the profession.

• On a 10-point scale, the top three reasons for considering a career change were: 7.6/10 — To achieve less stress.

7.5/10 — Achieve a better work/life balance.

7.1/10 — More meaningful/interesting work opportunities.

What continues to nag at me, since victory hasn’t been achieved yet, is that the data suggests that women seem to be less satisfied with the engineering profession than men and are, therefore, more prone to leave engineering than their male counterparts, particularly in mid-career.

Honestly, I am not bothered by anyone who changes their mind about their career path and decides to do something else. You only live once, and with a college degree (or two) under one’s belt, other career possibilities might be more enjoyable. If I wasn’t somewhat risk-averse and didn’t have a California-sized home mortgage, I might have decided to try to be a historian or a geologist exploring the Utah backcountry. To me, the underlying theme in this data is that, for many, structural engineering is or is becoming something to be endured and not a profession to be enjoyed.

I have previously noted that a significant factor in career satisfaction comes from being involved with people you like and working in an environment that gets as far away from drudgery as possible. Designing the same thing, day in and day out, or working on the same type of project

over and over would be hard for me to enjoy, just as would working under management that creates barriers to advancement and personal achievement or displays signs of favoritism or bias.

Another important source of career satisfaction and advancement for many people (which leads to promotions, better pay, more challenging roles, etc.) is engagement in business marketing activities, professional activities (like NCSEA committees), and volunteerism outside of the day-to-day in-house activities of the engineering office. These can be important and rewarding aspects of any career that create variety and keep things fresh.

However, these career-advancing activities, whether they present themselves in the normal course of the office, or are pursued outside the office, require support for the individual, both at work and home. This need for support is universal, but I sense that the availability of support may need to be even higher for women than men and that the lack of sufficient support may be partly the reason that so many people, particularly women engineers, have considered leaving the profession.

At the office, this means the demands on the individual may need to be lightened or re-arranged from time to time to allow time for other activities since one would assume that successful managers and companies want successful employees. When I went to business school, my manager, Loring Wyllie at Degenkolb Engineers, always made sure I was back in San Francisco for class starting at 6 pm, no matter what. Many of my classmates were not so fortunate. While true professionals must be willing to expend some of their own time in these pursuits, the employer shouldn’t make it more complicated than it needs to be or make the employee feel like they aren’t doing their fair share. Female and male employees should have an equal opportunity for extracurricular activities too. A manager that assumes that the woman is likely already overburdened or unavailable

is playing to the past stereotype.

In a two-working adult household, this means one person sometimes makes an extra effort to help the other so that careers aren’t negatively impacted. Just as we have welcomed women in the engineering office as the norm, we need to ensure it is being recognized as normal from a home perspective too. Having a two-working adult household can be demanding at times. Still, I would advocate for as much shared responsibility as possible to support women if we, as a profession, are to achieve equality. Let’s put the 1950s and the role models of the past behind us.

The woman that isn’t available for work travel, declines a promotion, can’t attend the evening marketing or business function, or the overnight business trip because her greater family demands is not being supported as she should be. Speaking as a dad, sharing the duty of cooking, cleaning, and getting the kids off to school or summer camp isn’t that hard, and if it is, try harder.

The goal should be equality of opportunity and letting the chips fall where they may. I am solidly against barriers erected, intentionally or not, that hinder anyone’s life, liberty, and pursuit of happiness, to quote Thomas Jefferson. There are many examples of extremely successful women in engineering, but we need more. I am not an “it takes a village” kind of person in that I believe that people can generally succeed on their own if they are committed to a goal and put in the effort. But everyone also needs to reconsider their own actions and take opportunities to chip away at or better yet remove barriers to make life better for others.■ John A. Dal Pino is a Principal with Claremont Engineers, Inc. in Oakland, California. He serves as the Chair of the STRUCTURE Editorial Board (jdalpino@claremontengineers.com).

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Drill Screws

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Prescriptive Design of Damped Moment Frames Simplifying building design with fluid viscous dampers by using an ICC-approved prescriptive method.

By Nathan Canney, Ph. D., P. E., Alan Klembczyk, Konrad Eriksen

F

luid Viscous Dampers (FVDs), or simply “dampers,” effectively dissipate seismic energy to reduce steel and foundation costs in new buildings and improve structural resilience, especially for critical facilities. A new building design procedure using dampers has been developed, validated, and published with certification from the International Code Council. This new design procedure opens the door for more damper applications in new steel structures by eliminating the requirements

12 STRUCTURE magazine

for nonlinear response spectrum analysis and peer review in damped buildings.

Introduction to Dampers The lateral force resisting systems (LFRSs) used as standard practice by structural engineers rely predominately on hysteretic damping

to dissipate seismic energy. This means axial deformation of braces, beam sections hinging in moment frames, rebar yielding, and concrete crushing in shear walls. For low-level shaking, the damage incurred can be negligible in most modern buildings. In moderate to large earthquakes, however, the damage can be significant, and while these buildings may meet the code intent of not collapsing, they can be too expensive to repair and result in a total loss. This became clear after the 2010/2011 Canterbury Earthquake Sequence in Christchurch, New Zealand. A study after the earthquakes found that over 65% of the “significant” buildings were demolished after the earthquakes, having lasting effects on the city (Gonzalez, Stephens, Toma, Elwood, & Dowdell, 2021).

FVDs have been successfully used in buildings for Figure 1 Example of a distributed damper configuration along the perimeter of a steel building. nearly 30 years to dissipate seismic energy, primarily through converting energy to heat rather than yielding structural and efficiency of the energy dissipated (Figure 3). elements. Typically, dampers are placed within building stories The fundamental equation that describes the behavior of FVDs is: to capitalize on the differential movement between floors (see Figure 1) and are often left exposed after the building is finalized F = CVα (Eq. 1) (see Figure 2).

The piston head of a damper contains specialized orifices, where -where F is the output damper force, C is the damping conthe size, quantity, and shape of these orifices control the damper stant, V is the induced velocity, and α is the damping exponent. force-velocity relationship and ultimately impact the amount For seismic applications, an α of 0.4 provides significant energy

(a)

(b)

(c)

(d)

Figure 2 Example Buildings with FVDs (a) 12 Moorhouse Ave., Christchurch, NZ; (b) Damper with integral extender at 350 California, San Francisco, CA; (c) Damper in coffee house at 555 Capitol Mall, Sacramento, CA; (d) Criterion Promenade, Santa Monica, CA.

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Figure 3 Fluid viscous damper components.

dissipation but with a lower output force when compared with linear dampers (α equal to 1.0).

FVDs have some unique properties which make them advantageous for building applications: • Reduce both drift and acceleration: Most LFRSs use increased stiffness to control drift but at the expense of higher floor accelerations. FVDs are unique because they help control inter-story drift while reducing floor accelerations.

• Does not impact building period: FVDs do not introduce stiffness into the structure (they have no displacementdependent characteristics) and, therefore, do not change the fundamental lateral period of the structure. This means that damped structures can be designed to have longer periods than their undamped counterparts, reducing their seismic response to earthquakes and the structure’s demands.

• Flexibility in placement: Because dampers do not introduce stiffness, they are not bound by the vertical regularity expectations of traditional LFRSs. Therefore, they do not require vertical stacking up the height of a building. Damper locations can shift from floor to floor depending on the functional and aesthetic needs, which often are heavily influenced by architectural preference.

• Out-of-Phase Behavior: FVDs are velocity-dependent devices, which means that their peak force occurs out-ofphase with the displacement-based seismic demands on the LFRS. When the building is at peak displacement, and therefore peak strain and stress in the LFRS, the dampers are experiencing zero velocity and, therefore, are not generating an output force. Conversely, when the building moves through the origin, the strain in the LFRS is zero, whereas the dampers generate their peak force corresponding to the peak velocity.

Damper Applications in Buildings Currently, most FVD projects in the U.S. are retrofits of PreNorthridge Steel Moment Frames and Nonductile Concrete Moment Frames in high seismic regions. For retrofits, reducing drifts without stiffening the building helps avoid costly foundation retrofits and can reduce demands on existing elements to within their current capacity.

14 STRUCTURE magazine

However, using FVDs in new construction has remained relatively limited to essential facilities like hospitals and emergency centers or applications where resiliency is a key objective. The code-driven requirements for NLRHA and peer review when designing with dampers are a barrier for many projects. To help mitigate these barriers, a new prescriptive method for designing new steel moment frame structures with FVDs was developed

by Taylor Devices, trademarked as the Taylor Damped Moment Frame™ (TDMF™).

This special damped moment frame system was developed and validated through the rigorous AC494 and FEMA P-695 processes and is officially approved in the International Code Council Evaluation Services ESR-4769 (https://icc-es.org/reportlisting/esr-4769/). This process included the design of over one hundred archetype structures and analyzing each structure’s collapse probability using advanced nonlinear analysis with a suite of 44 horizontal ground motion records. Through this process, it was demonstrated that structures designed in alignment with this design procedure meet the intention of the code, which is to have a less than 10% probability of collapse at an MCE-level seismic event.

The system design procedure utilizes Modal Response Spectrum Analysis (MRSA) for the steel moment frame analysis and design with modifications to ASCE 7 Chapter 12 to account for the positive impact of the dampers. The dampers are omitted from the analysis model and are designed separately through a prescriptive approach.

System Description The LFRS for the system is the steel special moment frame (SMF), where ordinary and intermediate steel moment frames are not eligible due to a lack of seismic detailing requirements.

The FVDs are supplemental to the LFRS. Therefore, they may be placed within the SMF or in so-called gravity frames. The frames that host the FVDs are designated as the Damper Frames (DFs). The SMFs and DFs are procedurally addressed separately, but in cases where the SMF and the DF share common elements (as is the case when the dampers are within the SMF), there is an iterative step where the damper force is introduced as a static demand back into the SMF. It should be noted that when the

FVDs are within a gravity frame, the beam-column connections do not have to be modified; they can remain as traditional gravity frame connections (see ESR-4769 for more detail).

Other system limitations include: • Flexible floor diaphragms as defined by ASCE 7 §12.3.1.1 or §12.3.1.3 are not permitted.

• Buildings must not have horizontal irregularity Type 1b, extreme torsional irregularity, as defined in ASCE 7 Table 12.3-1.

• A building height limit of 300 feet (measured from the base to the highest floor).

• At least two dampers must be installed in both principal directions on every floor for a minimum of four dampers per floor.

The dampers must be arranged with at least one damper on either side of the center of stiffness. While asymmetry in the damper placement is permitted, any damper-induced torsion must be accounted for through modification of the accidental torsion consideration (ASCE 7 § 12.8.4.2).

• Taylor Devices, Inc. must provide the dampers.

While the LFRS must be SMFs, there are no limitations on the rest of the structure so long as it meets the requirements above.

This means that the system is permissible with steel-framed buildings. Still, it can also be applied with mass timber, concrete, or masonry structures with an SMF lateral system and semi-rigid or rigid diaphragms. Notably, the ICC-approved system and prescriptive procedure do not apply to existing buildings.

Moment Frame Design The SMF is designed per AISC 341, 358, and 360 with demands from a modified MRSA. The system procedure is compatible with any pre-qualified or approved proprietary moment frame system.

The key design parameters and modifications to ASCE 7 Chapter 12 are: • R = 8, Ω0 = 3 (unchanged from Table 12.2-1) • Cd = 4.5 (instead of 5.5 from Table 12.2-1) • Scaling of forces: MRSA base shear is scaled to 75% of the ELF base shear, as opposed to 100% (modification to ASCE 7 § 12.9.1.4.1) • Scaling of drifts: When scaling MRSA drifts in accordance with ASCE 7 § 12.9.1.4.2, the factor CS used in Equation 12.8-6 is replaced with CS,d calculated as: CS,d = 0.35SD1 1R ≤ 0.5S1 1R ^I h ^I h e

(Eq. 2)

e

There is a 25% reduction in seismic base shear for force-controlled frames. Whereas for drift-controlled frames, there is an 18% reduction in floor displacements and a reduced minimum when Equation 12.8-6 from ASCE 7 controls. In both cases, the reduction in response leads to softer SMFs compared to undamped frames. Therefore, with longer periods and further down on the response spectrum, seismic demands are even further reduced.

Damper Frame Design The DF consists of the FVDs, extender braces, connections, gusset plates, beams, columns, diaphragms, and foundations

that form the damper force load path. Selection of the damper properties, C and α from Equation 1, comes from a prescriptive approach that occurs outside the MRSA used to design the SMFs. The damping exponent, α, is fixed at 0.4 for this system. The damping constant, C, is calculated using stiffness proportional damping up the structure’s height through the following series of equations.

First, a target linear damping constant, Cji(L), is calculated.

Cji(L) = b v

/

ki T1 1 2 r cos 2 i ji { cos ji j=1

(Eq.3)

ni

-where, βv = Target Viscous Damping Ratio = 0.25 θji = angle of inclination of the jth damper on the ith story measured from the horizontal φji = angle of the DF containing the jth damper on the ith story measured from the principal direction (zero degrees when the DF aligns with the principal direction) T1 = fundamental translational elastic period in the principal direction of interest ki = the ith story linear stiffness in the principal direction of interest n i = total quantity of dampers being considered in the principal direction of interest at the ith story If all the DFs are aligned along the principal directions, the terms ∑j=1nicos2 φji simplifies to ni. Equation 3 will produce a design where all the dampers on the ith floor in the principal direction of interest will have the same damping constant. For example, there are cases where varying damping constants are useful to balance an asymmetric damper placement. For this purpose, an alternative formulation of Equation 3 is provided in ESR-4769 which accounts for the summative damping, thereby allowing for varied damping constants on the same floor for a given principal direction.

The linear damping constant is translated to a nonlinear damping constant, Cji(NL), via equivalent energy dissipation at the peak pseudo-velocity. Individual damper displacements from the MRSA determine a pseudo-velocity for each device. At a preliminary analysis stage, the average MRSA drift could be used.

The pseudo-velocity is calculated as:

2r vji = T dji BF 1

(Eq. 4)

-where, dji = displacement (stroke) of the jth damper on the ith story measured from the MRSA analysis BF = base shear correction factor, which removes the effects of the base shear scaling from the DE level damper displacement = min (1.0, Vt /(CS,dW) Vt = modal base shear from ASCE 7 Section 12.9.1.4.2 W = seismic weight of the full structure The nonlinear damping constant is then calculated as: Cji(NL) = Cji(L) r (vji)(1−α) m

(Eq. 5)

-where λ = 3.582 for velocity [in./sec] and α equal to 0.4 The specified nonlinear damping constant, Cji(NL)spec, must be within a range of minus 10% and plus 30% of the values calculated N OVE M B ER 2023

15

4-Story Hospital

8-Story Hospital

SMF

BRBF

TDMFTM

SMF

BRBF

TDMFTM

Peak Inter-story Drift Ratio

0.93%

0.54%

0.82%

0.94%

0.65%

0.72%

Peak Roof Acceleration

0.77g

0.97g

0.60g

0.71g

0.92g

0.37g

Expected Annual Loss

$38,741

$51,827

$26,119

$45,926

$60,219

$19,068

Recovery Time–Re-Occupancy

4 mo.

8.5 mo.

2.4 wks.

4 mo.

6.1 mo.

8 days

Recovery Time–Functional Recovery

9.1 mo.

11 mo.

7.6 mo.

8.3 mo.

9.1 mo.

5.2 mo.

Table 1 Resiliency metrics comparison between archetypical hospital structures designed with alternative LFRSs.

by Equation 5, giving the designer flexibility and often allowing for smoothing of the damper designs over several floors.

The damper force and required stroke are then calculated to guide the engineer in selecting the damper size. The damper force at the DE level is determined based on Equation 1, except that the pseudo-velocity calculated in Equation 4 is modified to account for higher mode effects and nonlinearity. The DE level damper force is amplified to MCE level forces (×1.18) to correspond to the damper force ratings and by an overstrength factor (×1.66) to design the DF components.

The damper stroke is also determined with an “overstrength” factor, whereby the damper strokes determined from the MRSA are amplified by the importance factor, Ie, and a factor dependent upon the building’s Seismic Design Category and the number of stories. This amplification factor ranges from 2.5 to 3.5, reducing the dampers’ possibility of exhausting their provided stroke.

Additional requirements not covered here but provided in ESR-4769 account for a minimum extender stiffness, orthogonal damper effects, combined load effects between the lateral system and damper force, and torsional impacts of asymmetric damper placement.

Improved Performance and Resiliency with the TDMF™ System While the system was developed primarily to produce codecompliant buildings, namely structures that meet the collapse risk intention of the code, the use of dampers inherently improves performance and resiliency over traditional undamped counterparts. Dampers attenuate both story drifts and floor accelerations, reducing damage to structural and nonstructural components. This ultimately leads to less financial losses and building downtime following a seismic event.

Six archetypical hospital structures were designed to compare the resiliency of the system: 4- and 8- story buildings with LFRS of SMFs, BRBFs, and the system. These buildings were designed by a California structural engineering firm commissioned by Taylor Devices. FEMA P-58 and ATC-138 methodologies were employed to study resiliency using the software platform SP3 developed by the Haselton Baker Risk Group. Table 1 provides the results summary of each building’s performance, loss, and recovery time at the DE level.

These results demonstrate the system’s ability to, out of the box, contribute significantly to improving building resiliency. Floor accelerations and expected mean annual losses are significantly reduced. Time to achieve re-occupancy, a crucial criterion for 16 STRUCTURE magazine

any hospital, shifts from months to days. Further, the changes needed to make the system achieve immediate occupancy or higher resiliency standards are minor compared to the other systems and rest largely on decisions outside the lateral system.

The dampers are designed to maintain their full function even after experiencing an MCE-level seismic event and do not require replacement or servicing.

With our industry increasingly acknowledging the reality of our code minimum building performance, the system will become a critical tool in the toolbox to provide code-level buildings that increase our community resiliency. The system design procedure opens the door for more applications of dampers in new steel structures, removing barriers of NLRHA and peer review, which add cost and time to any project. Additionally, removing NLRHA increases the accessibility of dampers to many engineering firms that might otherwise not have considered a damped design. This new system prescriptive design approach can be a key component in the structural engineering community’s efforts to improve public safety by providing facilities that do more than just survive earthquakes without collapsing – we can help make our cities resilient to avoid the post-earthquake losses seen elsewhere.■

Full references are included in the online version of the article at STRUCTUREmag.org.

Nathan Canney, Ph. D., P. E. (nathancanney@taylordevices.com) is the Director of Structural Engineering at Taylor Devices, Inc. (TDI). Prior to joining TDI in 2020, Nathan worked as a structural engineer and was faculty in the Department of Civil and Environmental Engineering at Seattle University from 2013 to 2017.

Alan Klembczyk (alanklembczyk@taylordevices.com), President and Board Member of Taylor Devices, Inc. since 2018, has spent 35 years with the company. After obtaining a degree in Mechanical Engineering, much of his career has involved designing and developing shock and vibration mitigating products to improve performance during transient events including earthquakes.

Konrad Eriksen (konraderiksen@taylordevices.com) is the Structural Products Sales Director at Taylor Devices, Inc. since 2020. He has more than 30 years of experience manufacturing, developing, and selling highperformance construction devices that include dampers and base isolators.

He also has 10 years of hands-on commercial construction experience in New Zealand.

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structural DESIGN Design of Stainless Steel Bolted Connections in Accordance With the Recently Published ANSI/AISC 370

Design rules for bearing-type and slip-critical bolted connections made of stainless steel are now available.

By Francisco Meza, Ph.D. and Nancy Baddoo

I

n the December 2022 edition of Structure Magazine, Benjamin Baer gave an overview of the new AISC design specification for structural stainless steel, ANSI/AISC 370, and introduced the provisions for designing members. This article focuses on the design of bolted connections in structural stainless steel.

Introduction Stainless steel bolts are used in a wide range of architectural and industrial applications (such as water treatment, food processing, chemical processing, marine structures, and nuclear facilities) where carbon steel bolts would be simply inadequate. They are availFigure 1 Typical detail for connecting dissimilar materials to avoid galvanic corrosion for bolts installed to the able in various alloys and conditions, offering snug-tight condition in a water-shedding service environment. different strengths, corrosion resistance levels and the ability to operate at high and low temperatures. Even though stainless steel bolts have a higher initial group providing unique properties and a range of different corrosion cost than their carbon steel equivalents, the savings from a long levels. ANSI/AISC 370 covers the design of members and bolts made life with low maintenance can easily outweigh these initial costs in from the austenitic and duplex stainless steel alloy groups, which are demanding applications. As well as being essential for connecting the most commonly used structural stainless steels. The design of stainless steel members, stainless steel bolts are also suitable for con- precipitation hardening stainless steel alloys is also covered for tennecting galvanized steel and aluminum members. They are also used sion members, fittings and fasteners where higher strength is needed. in timber construction because they are resistant to chemicals used A summary of the design rules and key recommendations for bolted to treat timber, and also resistant to corrosion caused by acetic acid connections is given below. which is emitted by timber.

Design of Bolted Connections The design of stainless steel bolted connections in the US has become significantly easier thanks to the release in 2021 of the Specification for Structural Stainless Steel Buildings (ANSI/AISC 370) and its accompanying Code of Standard Practice for Structural Stainless Steel Buildings (AISC 313). The second edition of Design Guide 27 Structural Stainless Steel (AISC DG 27), released in 2022, also gives practical guidelines for designing stainless steel bolted connections and provides design tables for determining their strength.

Stainless steels can be classified into five basic groups, with each 18 STRUCTURE magazine

Material Selection and Specification An important consideration when selecting a stainless steel bolt, nut, and washer is to ensure they have a corrosion resistance at least as good as that of the most corrosion-resistant material being joined. When a structure is exposed to a corrosive environment, the connections are the parts most affected by the aggressive environment as they are susceptible to crevice corrosion, stress corrosion cracking, and galvanic corrosion if dissimilar metals are connected.

The potential for crevice corrosion at a connection should be considered in any environment where moisture is present and surface deposits may accumulate, particularly those containing chloride salts.

Bracing connection in a stainless steel canopy at Porto Airport, Portugal. Photo on right courtesy of TriPyramid Structures, Inc.

Crevice corrosion can occur at locations that are shielded or occluded and have insufficient localized oxygen concentrations that are too low to maintain a passive film. Possible sources of moisture include spray, ponding, rain, fog, condensation, and humidity. When crevice corrosion is a concern, the most common design options are to seal the joint, modify the design to reduce the risk of regular moisture accumulation or select a stainless steel alloy with sufficient crevice corrosion resistance for the service environment.

The development of chloride stress corrosion cracking (SCC) requires the simultaneous presence of tensile stresses and a chloride-rich environment, which is unlikely to be encountered in standard building atmospheres. Service environments potentially prone to chloride SCC include indoor swimming pool areas, specific industrial and food processing environments, and structural members in areas with high coastal or de-icing salts accumulation. The stresses do not need to be very high in relation to the material's yield stress. They may be due to loading or residual effects from manufacturing processes.

Caution should be exercised when stainless steel members containing high residual stresses (e.g., due to cold working) are used in chloriderich environments. Duplex stainless steel bolts usually have superior resistance to SCC than the austenitic stainless steel bolts covered in ANSI/AISC 370. Precipitation hardening stainless steel bolts should not be used when SCC is a concern. The potential for galvanic corrosion should also be considered in bolted connections that combine stainless steel with other metals. In this type of connection, the dissimilar metals must be isolated from one another electrically. This usually includes using insulating washers and bushings to prevent contact between the bolt and members being joined so that galvanic corrosion is avoided if there is moisture infiltration or condensation within the joint. Figure 1 shows an example of galvanic separation of dissimilar metals for a bearing-type connection, in which the bolt is snug-tightened, in a water-shedding application. The connection in Figure 1 would not be an appropriate design for a regularly or continuously immersed condition or where crevice corrosion of the stainless steel was a concern. Because of the need for isolating material,

slip critical connections cannot be used for joining dissimilar metals.

Several ASTM standards cover stainless steel bolts. Most stainless steel bolt standards have a corresponding standard for the nut unless the nut is specified in the same standard as the bolt (this is the case for ASTM A1082/A1082M, which covers stainless steels suitable for bolts and nuts).

The ASTM standards covering stainless steel bolts are given in Table 1. There are no ASTM standards for stainless steel washers.

However, requirements are given in ANSI/AISC 370 regarding material selection and hardness.

For the austenitic stainless steel bolts, although all the standards give the same values for yield strength and tensile strength of bolts in the annealed condition (Fyb = 30ksi, Fub = 75 ksi), the strengths in the cold worked condition and strain-hardened condition depend on both the alloy and the diameter of the bolt. Strength and size are inversely correlated because it is difficult to achieve uniform work hardening through the thickness of a bolt. For the duplex stainless Material

Austenitic

Duplex

Bolt

Nut

Common alloys

ASTM F593

ASTM F594

Group 1 (304, 304L), Group 2 (316, 316L)

ASTM A320

ASTM A962/ A962M

B8 (304), B8M (316)

ASTM A193/ A193M

ASTM A194/ A194M

B8 (304), B8M (316)

ASTM A1082/ ASTM A1082/ 2205 A1082M A1082M

ASTM F593 ASTM F5942M 630 (17-4) Precipitation ASTM A1082/ ASTM A1082/ Hardening 630 (17-4) A1082M A1082M Table 1 ASTM product standards for stainless steel bolts and nuts.

N OVE M B ER 2023

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steel bolts, the yield and tensile strengths vary between alloys, and in the case of precipitation hardening stainless steel bolts, they also vary with the heat treatment. (ANSI/AISC 370 permits the precipitation hardening alloys only in heat treatment condition H1150 to reduce the risk of hydrogen embrittlement and SCC).

Surface Class

Slip coefficient, µ[a]

SSB

0.20

SSC

0.40

SSD

0.50

Values include the potential loss of preloading force due to timedependent relaxation.

[a]

Structural Performance of Stainless Steel Bolted Connections Stainless steel bolts behave differently from carbon steel bolts due to the different material composition and non-linear stressstrain characteristics. This means it is only possible to adopt some of the design rules and installation procedures used for carbon steel bolts.

Bearing-type Connections As with carbon steel, the strength of stainless steel bearing-type connections may be governed by the failure of the bolt itself or the connected plates. Failure of the bolt may be due to shear, tension, or a combination thereof, while the connected plates may fail in bearing or tearing when the spacing between bolts or the spacing between the bolt and the edges of the plate is not sufficient.

The minimum bolt spacing and edge distance requirements are the same for stainless steel and carbon steel bolted connections.

The strength provisions given in ANSI/AISC 370 for stainless steel bolts are similar to those for a carbon steel bolt of the same diameter and tensile strength (Fub). As for carbon steel bolts, the shear and tensile strengths of a stainless steel bolts are calculated by multiplying the nominal (tensile or shear) stress by the nominal unthreaded area of the bolt. If the bolt is subjected to combined tension and shear, the strength is predicted using the same interaction equation used for carbon steel bolts. However, due to the wide variety of stainless steel bolts on the market, each with different tensile strength (Fub), rather than tabulating the nominal (tensile and shear) stresses, ANSI/AISC 370 provides simple expressions for calculating these values.

For stainless steel bolts in tension, the nominal tensile stress is given by 0.75Fub, where the 0.75 coefficient accounts for the reduced cross-sectional area of the threaded part of the bolt.

Similarly, if shear failure occurs in the threaded portion of the bolt, the nominal shear stress of the unthreaded portion of the bolt (0.55Fub) is reduced to 0.45Fub to compensate for the reduced cross-sectional area of the threaded part of the bolt. The 0.55 and 0.45 coefficients include a 10% reduction in strength to account for the detrimental effect of the uneven redistribution of stresses in long-bolted connections.

The resistance factor used in calculating the available strength of austenitic and duplex stainless steel bolts is the same as that used for carbon steel bolts. However, it was reduced by 10% for precipitation hardening stainless steel bolts to account for the lack of sufficient test data on this type of bolt.

The equations for calculating the bearing and tearing strength of stainless steel plates have a similar form to those used for carbon steel plates, with the tearing strength increasing proportionally to the bolt-to-bolt or bolt-to-edge distance until the bearing strength is reached. However, stainless steel plates will exhibit lower bearing and tear-out strengths due to the gradual softening of the material. As for carbon steel bearing-type connections, 20 STRUCTURE magazine

Table 2 Slip coefficients for stainless steel slip-critical connections given in ANSI/ AISC 370.

the bearing/tear-out strength of the plate can be determined for the case in which the deformation at the bolt hole under service loads is or is not a design consideration.

Slip-critical Connections Slip-critical connections are required when deformations in bolted connections must be limited to predefined values under service or ultimate loads. Typical applications can be found in bridges, cranes, radio masts, and towers of wind turbines. In this type of connection, the shear force is transferred by friction between the faying surfaces, and therefore, ensuring that the level of pretension in the bolt and the slip coefficient of the faying surfaces are achieved is key to guaranteeing the strength of the connection.

ANSI/AISC 370 covers the design of slip-critical connections in which the bolt is made of austenitic or duplex stainless steel with a specified minimum yield strength F yb ranging from 80 ksi (550 MPa) to 130 ksi (900 MPa) and faying surfaces made of stainless steel plates. The reader should be aware that there is a typographical error in the upper bound limit of Fyb (<= 116 ksi (800 MPa)) given in the current version of ANSI/AISC 370, which will be corrected in the next edition in 2025. The use of slip-critical connections comprising stainless steel bolts and carbon steel plates is outside the scope of ANSI/AISC 370. This is because in this type of connection, the dissimilar metals would need to be isolated similarly to the connection shown in Figure 1 with the washers and bushings made of an incompressible material, for which there is currently no data.

To limit the surface pressures on the clamped package and thus minimize pretension losses due to plastification of the surface, ANSI/AISC 370 requires washers hardened to at least 290 Brinell HBW to be placed under both the bolt head and the nut.

Embedding of the surface can also be minimized using washers with a diameter larger than the nut or the head of the bolt.

Galling can occur when the bolt is highly pre-tensioned or torqued, and stainless steel bolting assemblies are more susceptible to this than carbon steel bolting assemblies. Therefore, the use of lubrication during the installation of slip-critical bolting assemblies is essential to minimize galling. It is recommended that the lubricant is applied between the paired threads and the bearing surfaces between the nut and the washer.

The nominal strength of stainless steel slip-critical connections is calculated by the same equation used for carbon steel slipcritical connections, which is given by Equation 1, where: Du is a multiplier that reflects the ratio of the mean installed bolt pretension to the specified minimum bolt pretension; Tb is the specified minimum bolt pretension, hf is a reduction factor that is used when two or more fillers are placed between the connected

Rz[b]

Rt[c]

Surface Class[a]

µin.

µm

µin.

µm

SSB

1400

≥ 35

2000

≥ 50

SSC

1800

≥ 45

2400

≥ 60

SSD

2200

≥ 55

2800

≥ 70

Surfaces must be grit-blasted.

Rz is the surface roughness according to ASTM D7127. [c] Rt is the surface roughness according to ASTM D4417. [a] [b]

Table 3 Definition of surface classes for stainless steel slip-critical connections given in ANSI/AISC 370.

parts (hf is the same as for carbon steel slip-critical connections); ns is the number of slip planes, and μ is the slip coefficient for the faying surfaces.

Rn μDuhfTbns

(1)

the defined roughness values given in Table 3. If a faying surface with different surface roughness is used, tests can be conducted according to the testing method given in AISC DG 27 to determine the appropriate slip coefficient to use in the design.

When a slip-critical connection is subject to both tension and shear, the tensile force reduces the clamping force between the faying surfaces, reducing the slip resistance of the connection. This effect is accounted for by applying the same reduction factor used for slip-critical connections made of carbon steel.

Summary Stainless steel bolts have a long track record of providing durable solutions for connecting a wide range of materials. The design process is similar to that of carbon steel bolts. Care is needed in specifying stainless steel bolts as there are a range of alloys and conditions (strengths). Guidance is available in ANSI/ AISC 370 and AISC DG 27. In the coming months, an RCSC/ AISC stainless steel bolting guide will be prepared, with essential information for specifiers and designers all in one place.■

TAYLOR DAMPED D E™ MOMENT FRAME SIMPLIFIED

Photo credit: John Doogan/WSP

taylor

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The resistance factors used to calculate the available slip resistance for different hole types (i.e., standard size, slotted and oversized holes) are also the same as those used with carbon steel slip-critical connections.

An important distinction when calculating the resistance of stainless steel slip-critical connections is that the minimum bolt Francisco Meza, Ph. D., is a principal engineer at the Steel Construction pretension Tb is taken as 70% of the yield strength of the bolt Institute. He has worked on a number of projects developing design multiplied by the net tensile area of the bolt. This definition of guidance and specifications for structural stainless steel, and has Tb differs from that for carbon steel slip-critical connections, undertaken a key role in drafting ANSI/AISC 370-21, Specification for which is based on 70% of the tensile strength of the bolt. The Structural Stainless Steel Buildings. He also co-authored the Second Edition reason Tb is based on the yield strength of the bolt is that tests of AISC Design Guide 27 Structural Stainless Steel. on stainless steel bolts have shown that it is only sometimes posNancy Baddoo is an Associate Director at the Steel Construction Institute. sible to guarantee a minimum bolt pretension as high as 70% of For the last 12 years, she has chaired the European technical committee the tensile strength. responsible for the stainless steel Eurocode, EN 1993-1-4. She wrote the To ensure that the specified minimum bolt pretension is achieved, First Edition of AISC Design Guide 27, co-authored the Second Edition, ANSI/AISC 370 requires that the stainless steel bolting assembly, and played a key role in drafting ANSI/AISC 370-21. as well as its installation process, meet the requirements of a Bolt Tightening Qualification Procedure, such as the one included in AISC DG 27, which covers the turn-of-nut, the calibrated wrench, and the combined methods.

The value of the multiplier Du in devices inc.

Equation (1) was conservatively set to 1.0 for stainless steel slip-critical connections because there was insufficient data to justify a higher value.

This contrasts with the value of 1.13 used for carbon steel slip-critical connections, for which significantly more data is available on the differDAMPER R DESIGN ent bolt installation methods.

In stainless steel slip-critical connections, the slip coefficient μ is primarily dependent on the roughNo N o Peer Review wR Required equ uired ness and asperity of the faying No Tim Time-History m e History A Analysis n a surfaces, the latter being conQuick Design Times trolled by the type of media used to blast the surfaces. Table 2 shows the current slip coefficient values given in ANSI/AISC 370. These values are only applicable to grit(716) 694-0800 | www.taylordevices.com ESR-4769 blasted faying surfaces that meet

N OVE M B ER 2023

21

Building west elevation overview looking at Lake Superior.

Vision Northland - Minnesota Hospital Design Braces for Winter and the Future

By Colleen F. Blackwell, P. E., Stephen G. Bartal, P. E.

E

ssentia Health’s Vision Northland project, located on the shores of Lake Superior in Duluth, Minnesota, includes approximately 928,000 square feet of new construction and roughly 120,000 square feet of renovation to existing facilities and is scheduled to be completed by the end of 2023. The project has created a civic asset capable of being both the standard bearer of healthcare in the region and a physical celebration of the community it serves. This formidable project is a healthcare destination for Northern Minnesota. While receiving the benefits of world-class healthcare, facility occupants can enjoy year-round expansive views of Lake Superior.

22 STRUCTURE magazine

New construction adds a nine-story Out-Patient building, a 17-story In-Patient tower to Essentia’s existing Duluth clinics, and a two-story over-build on existing Duluth Clinic DC-1.

The In-Patient tower reaches a height of 290 feet, topped out with a mechanical penthouse and helipad. The project had to overcome many complicated region-specific issues, including wind loads from Lake Superior, heavy snow loads with drifting, a frigid winter climate necessitating close attention to thermal continuity through the exterior building envelope, and the steep slopes of existing grade. The complexity of many of these issues

Figure 1 Overall North to South Building Section.

was compounded by the new project’s location being constructed change and the requirement to align with existing building floor adjacent and attached to an existing operating hospital. In addi- levels results in the varying floor-to-floor heights in the new tion, design and construction occurred during a global pandemic. structure of up to 24 feet.

The design team also had to work closely with the construction team to deliver early structural packages to meet the demanding construction schedule. Construction documents for foundations Site-Specific Geological Challenges and early steel mill order packages were delivered to the construction team four months and eight months before the overall From the very beginning of the project, the primary goal was to architectural design was finalized. provide a healthcare destination for the region. Achieving this goal The project is built into the hillside of Duluth, Minnesota, on required careful site selection for the new structure to integrate the shore of Lake Superior. The facility climbs the hillside start- the new building into the existing facility. Siting for the building ing from the shoreline of the lake at Superior Street, rising 100 feet in grade change up to 2nd Street.

The footprint covers approximately 600 feet, spanning an alley and frequently traveled 1st Street.

The street crossings create several long-span transfer conditions. In addition, the new facility is also built integral with many existing buildings across multiple city blocks. See Figure 1 for a building cross section.

Interconnecting the new building with existing buildings on the campus was complex due to the varying ages of the existing construction and the differing floor-to-floor heights. The new building is laid out to align in elevation at the north end with 2nd Street (Level 6 in the new building) and at the south end along Superior Street at Level 1. The site elevation 60 feet tall, 18’ thick cantilever retaining wall at north end of site.

N OVE M B ER 2023

23

V-Column support of Out-Patient Building during construction.

Braced Frame Connection to 35 feet tall, W14x605 column at north atrium.

required the demolition of several existing buildings and the consolidating of some existing facility services.

The project is located within what is known as the iron range of Northern Minnesota, which means this area is known for hard rock formations. A detailed geotechnical engineering evaluation

was performed to evaluate existing site conditions and develop criteria for the new foundations. Generally, the soils present at the site comprise fill-over glacial deposits, weathered bedrock, and bedrock. With maximum column loads of approximately 8,100 kips, it was necessary to construct all new foundations on existing bedrock. The in-situ rock was determined to have an allowable bearing capacity of 50 tons per square foot. Foundations for the new structures consist of spread footings bearing on the rock.

Rock-anchored mat foundations are utilized at the braced frame locations to resist uplift and overturning forces.

Depth to competent bedrock varied, and reaching bedrock for foundation bearing and removal for the new building program required extensive excavation, drilling, and blasting of the existing bedrock. Approximately 79,500 cubic yards of rock removal was required for the project, including approximately 53,000 cubic yards of the rock requiring blasting. With an open and operating hospital on an active construction site, rock blasting was limited to specific times of day to limit the impact of sound and vibration on existing hospital functionality. Excavation work and rock removal took four months for the Outpatient building and ten months for the Inpatient building.

Building Superstructure Overview Out-Patient overview of Southeast corner with view of brick cladding projections and offset columns.

24 STRUCTURE magazine

Due to location and labor constraints (need to describe the constraints), structural steel was selected as the material of choice for the main structural frame. The structural system for the

Out-Patient overview of Southwest corner with view of curtainwall setbacks and 15 foot cantilevered wingwall.

project is comprised of conventionally framed composite metal deck floor slabs. Steel framing is designed with headed studs to take advantage of the composite behavior and help reduce steel tonnage. The total project steel tonnage was approximately 12,000 tons).

The building utilizes steel-braced frames as the lateral forceresisting system. Since the project is located in a region of very low seismicity, the wind loads from the open lake surface and surrounding terrain controlled the lateral force-resisting system design.

Most of the steel for the project was typically ASTM A992 Grade 50. However, all truss members, plate girders, braced frame members, and large column sections utilize ASTM A913 Grade 65 steel to reduce steel sizes and keep the steel as economical as possible. At the time of design, grade 65 steel was only being rolled in column sections.

The Out-Patient building steel trusses span over 50 feet and support five stories above, allowing vehicle clearances in the alley.

Similarly, the inpatient building spans over highly trafficked 1st Street and an open ambulance bay area using story-height steel trusses (18´-0˝ deep) spanning up to 95 feet and supporting up to three stories above. Two pedestrian bridges also took advantage of long-span steel trusses over 80 feet in length to connect new and existing parking garages across 4th Avenue.

In-Patient west elevation showing transfer over 1st Street.

Wind tunnel results yielded wind loading significantly lower than those required by code. Per ASCE 7 chapter 31.4.3 Limitations on Loads, “Loads for the main wind force resisting system determined by wind tunnel testing shall be limited such that the overall principal loads in the x and y directions are not less than 80 percent of those that would be obtained from Part 1 of Chapter 27 or Part 1 of Chapter 28.”

The code describes a few conditions that could allow for further reduction; however, the site did not meet these criteria; therefore, the reduction in wind loading was limited to a twenty percent reduction from the code-prescribed loads.

Limiting structural vibration was also a primary focus in the design of the new structure. Expansive surgical and imaging floors with stringent limitations on allowable vibration were of particular concern in bays with long-span conditions and surgical suites over transfer trusses. An initial analysis by EwingCole’s design team verified that all structural steel met the criteria for sensitive equipment per American Institute of Steel Construction (AISC) Design Guide 11 – “Floor Vibrations Due to Human Activity.”

Development of Project Specific Loading Architectural programming combined with the site located in northern Minnesota created unique snow and wind loading conditions. The project was designed in accordance with the governing Minnesota state building code, which references IBC 2012 along with subcode ASCE 7-10. Code-prescribed wind loads for this site were developed based on an exposure category D due to the proximity to Lake Superior, even though the site is in an urban area. Despite the code provisions, the site terrain made the design team question how much wind loading this building would experience. To optimize the structural design, a wind tunnel study, and a snow and ice evaluation were performed

by Novus Environmental Inc. (now SLR). A wind tunnel study was performed for a few reasons: to refine or reduce code-required wind loads and to develop components and cladding loads.

In-Patient North Entrance Canopy and north façade with brick support at story mid height.

N OVE M B ER 2023

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process was used to refine the framing design based on actual room layout and excitation. The framing was analyzed based on walking excitation within the rooms and outside of the room in the corridor, as well as the impact of mechanical equipment in close proximity to sensitive areas.

See Figure 2 for a mapping of floor vibration performance at one of the typical patient floors.

Meeting Architectural Goals

Figure 2 In-Patient Tower floor plan vibration performance mapping.

EwingCole’s design team collaborated closely with consulting firm Novus Environmental to further refine the analysis.

Vibration limits were set per the Facility Guidelines Institute (FGI) 2018 Hospital guidelines reference. For example, Patient Care areas were designed for a Footfall Vibration RMS Velocity (micro-inches per second) of 6,000, general surgery at 4,000 or less, and imaging at 2,000 or less. Finite Element analysis was performed to develop a full understanding of framing vibration accelerations. With such stringent criteria, an iterative

Building west elevation overview looking north in the evening.

26 STRUCTURE magazine

The landmark nature of the project led to many monumental architectural features and open interior spaces. The outpatient building boasts a six-story atrium made possible using horizontal trusses on levels two through four to brace the exterior curtainwall and transfer wind loading at a multistory atrium space.

Clearstory curtainwall spanning up to 35 feet without intermediate support was constructed from Levels 4 to 6.

The north In-Patient Tower atrium also has a tall expanse of curtain wall. The north curtain wall is hung from the Level 8 steel to minimize vertical mullion sizes and maintain the clear without horizontal support.

The majority of the Out-patient Building’s southeast corner is supported by what is known as a “V” column at the Lake Superior entrance. This column supports nine levels of the Out-Patient building through transfers onto plated wide flange members, all fabricated to meet architecturally exposed structural steel (AESS Category 3) requirements. This category of AESS steel required a visual mock-up of the welds and painted finish for architectural approval. All welds were ground smooth, and the paint was finished with a smooth texture since the columns were visible from a close distance.

Along with the curtainwall, the building façade features many elements, including metal panels, Porcelanosa stone, sunshades, and projecting brickwork to blend the new with existing buildings. HSS tube hangers, braces, and horizontals support a complex brick façade. The exterior brick mimics the brickwork of neighboring 19th and early 20th-century buildings with projections. It reveals detailed hung and cantilevered steel, all supported by spandrel beams offset 13 feet from the closest columns. Detailed calculations were required to meet the brick live load deflection limits.

Steel plate girders were used on Level 6

of the Out-Patient building so that columns supporting floors 7 through nine could be pulled toward the interior of the building. This allowed ample outdoor balcony space on the sixth floor with no columns.

HSS tubes create multiple extended wing walls with curtainwall façade. These tubes featured end plate connections with thermal isolation material separating the external wing wall feature and interior conditioned space. Thermal break material was integrated into the cantilever connections to maintain envelope continuity.

South Out-Patient Atrium at project completion.

Designing for the Future This project used overbuild capacity in the existing 1st Street Duluth clinic design to add two stories. Additionally, structural provisions were made for an expanded future roof garden for women and children. The Out-Patient building was also designed to support a future expansion of two stories atop the building.

Novel projects require and promote innovative project solutions.

Essentia’s Vision Northland presented abundant and unique challenges based on architectural goals, existing conditions, underlying geology, site topography, and a harsh northern climate. The structural design was optimized for long horizontal and vertical spans in the presence of refined yet substantial wind and snow loads. On-going hospital operations in existing buildings required a construction schedule to minimize the effects of blast vibrations. In addition, the new structure was designed for stringent occupant footfall vibrations. Growth of the new structure was considered with future overbuild loads in the current iteration of the design.

On July 20, 2023, Essentia Health held the Grand Opening for the hospital to welcome patients and caregivers with all the accolades one can imagine for such a formidable facility. The exhaustive efforts of the design team, contractor, and administration led to a successfully designed, detailed, fabricated, and erected facility to provide Duluth, Minnesota, with a state-of-the-art healthcare facility ready for this generation and the next.■

Project Team Owner: Essentia Health Architect: EwingCole Structural Engineer: EwingCole Acoustic/ Vibration Consultant: Novus Environmental Inc. Construction Manager/Concrete Contractor: McGough Construction Steel Fabricator: LeJuene Steel Co. Steel Erector: Danny’s Construction Co.

Colleen F. Blackwell, P. E. is a Principal with EwingCole in Philadelphia, Pennsylvania. Member NCSEA, SEAOP and the Delaware Valley Association of Structural Engineers (DVASE). (cblackwell@ewingcole.com) Stephen G. Bartal, P. E. is a Project Engineer with EwingCole in Philadelphia, Pennsylvania. (sbartal@ewingcole.com) North Atrium with 35 foot tall braced frames accented by architectural cladding.

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Fully constructed Terminus. Photo courtesy of TBC.

Terminus

Exploring new frontiers in lateral systems for mass timber.

By Ilana Danzig, P. E., S. E., P. Eng, Struct Eng, M. Eng

T

he completion of a 5-story mass timber commercial office building named Terminus in the city of Langford, British Columbia, on Canada’s west coast, marks an exciting first in the world of seismic systems for timber structures. The building includes 4 stories of postand-beam mass timber with Douglas Fir glulam beams and columns and Spruce-Pine-Fir (SPF) cross-laminated timber (CLT), all supplied

by Structurlam. While the structure and its detailing are clean, elegant, and modern, what is truly unique about Terminus is its steel-timber hybrid seismic force resisting system (SFRS).

Terminus was developed to house the head office of Design Build Services (DBS), the building’s owner and the design-builder behind many Langford developments. As an early adopter of mass timber, DBS embraced the material fully, recognizing its innate benefits like the potential for fast speed of construction, low structural embodied carbon, and a beautiful, finished look. For Terminus, DBS sought out a building that had the architectural polish of modern mass timber without defaulting to a steel or concrete SFRS. The site is one of the highest-seismic regions of Canada; with significant lateral demands on the project, the lateral load-resisting system was a critical element

28 STRUCTURE magazine

of the building design. Aspect Structural Engineers, familiar with a large assortment of options for mass timber lateral systems, and with a mandate to spend the time and effort to make the right choice for the project, undertook a deep dive into investigating SFRS options that were high-performing, complementary to the exposed timber structure, and cost competitive.

Glulam Chevron Braces One of the simplest mass timber SFRS options is glulam braced frames, which are permitted in the National Building Code of Canada (NBCC) with a combined RdRo factor of either 2 or 3, depending on the ductility of the detailing. RdRo is approximately equivalent to the Seismic Response Modification factor, R in ASCE 7, and fulfils the same function.

Neither the NBCC nor the Canadian timber code, Engineering Design in Wood (CAN/CSA-O86), include specific detailing provisions for timber braced frames, and it is left to engineering judgement

Preferred connection yield modes. Image courtesy of Aspect Structural Engineers.

Rendering of a ductile connection with tight-fit pins and multiple knife plates. Image courtesy of Aspect Structural Engineers

by experienced engineers to demonstrate through research, testing, and project precedents that the system is adequately ductile. Best practices, research, and precedents indicate that the timber elements shall remain elastic and largely damage-free in a seismic event, while the connections between members (typically steel elements) provide the system’s ductility. As in many other seismic systems, steel elements that can undergo substantial and predictable plastic deformation without failure make good options for connectors in mass timber braces.

Tight-fit pin connections, with many small diameter pins that are tightly fit into the timber member and steel knife plate, can ensure a yield mode with bending of the pins and some wood crushing rather than splitting or fracture of the timber as long as limits are imposed on the pin spacing and slenderness. Tight-fit pin connections have been shown to provide adequate ductility for an RdRo of 2 or 3, however, this is still a relatively low response modification factor relative to other SFRSs and would make for a very high seismic demand building-wide. A low RdRo factor can also drive up elements that are capacity protected (designed to exceed the expected capacity of the SFRS), including timber member sizes, foundations, and the diaphragm.

damping and energy dissipation stemming from the yielding of steel elements, these dampers function by clamping steel plates together and providing damping via the friction generated when the plates slide past one another. A major advantage of these devices is that there is no sacrificial structural component; energy is dissipated via friction instead of steel yielding. These are typically used in steel and concrete structures, both in seismic retrofits and new construction.

There is research demonstrating that the use of these dampers provides a hysteretic performance approximately on par with Buckling Restrained Braces (BRBs), providing a pathway to an equivalent response modification factor.

One disadvantage of these dampers is the lack of self-centering capability of the system. Because the dampers use friction to dissipate energy, there is no elastic component to the deformation that can pull the system back to center once the slip friction is exceeded. Combining slip-friction braces with elastic supplementary moment frames is often recommended to provide nominal elastic centering to the building. Mass timber moment frames are not unattainable, however in analyzing the building and determining the stiffness that would be required for the moment frames to be effective, it was clear that the mass timber elements within the moment frames would need to be significantly upsized, increasing the overall project cost.

Slip Friction Damper Intrigued by slip friction dampers supplied by Quaketek, a Quebec-based manufacturer, and their capacity to dissipate a large amount of seismic energy, the team investigated their use for the Terminus project. Instead of

Photo of the building with braces exposed. Photo courtesy of Aspect Structural Engineers image, copyright Dasha Armstrong.

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approach instead of an equivalent static or linear dynamic analysis that is more typical of Canadian and US seismic design. ASCE 41 is an excellent resource for performance-based design of lateral systems that do not have an established response modification factor; the performance level (drift limits, in this case) is established first, and the device is tuned to achieve those limits with adequate damping in order to limit the force demands.

Buckling Restrained Braces (BRBs) BRBs are steel braces that have a yielding steel core that is restrained from buckling by a grout-filled outer steel casing. These braces are increasingly used in mass timber structures in high-seismic regions due to the well-established code procedures for design and testing, and the high response modification factor (RdRo = 4.8 in Canada and R = 8 in the United States). In a mass timber building, BRBs will typically necessitate steel columns and beams at each braced bay to fully comply with prescriptive code requirements.

Since the client was interested in the appearance of the mass timber structure and keeping as much of it exposed as possible, a novel approach to BRBs was developed for this project: using BRB braces within an all-timber frame instead of substituting the glulam columns and beams for steel within the braced bays. Finding no precedents for this approach, the team looked to first principles, the intent behind the BRB Canadian code provisions, and timber best practices to ensure compatibility between the timber and steel elements and maintain overall performance of the BRB braced bays.

Photo of a column base at brace with tight fit pin connections. Photo courtesy of Aspect Structural Engineers.

Tectonus Resilient Slip Friction Joints (RSFJ) Tectonus RSFJs are devices that have been developed for timber structures in for high-seismic regions. This system was invented in New Zealand in the aftermath of the 2010 Christchurch earthquake and the resulting catastrophic damage that occurred. RSFJs are springloaded slip friction dampers, so there is a built-in self-centering component to the system. Rather than reduce the building’s base shear via energy dissipation due to friction alone, these devices produce a flag-shaped hysteresis, reducing the building’s base shear by providing energy dissipation and by increasing the period of the building while ensuring that the drift limit is maintained. The RSFJs can result in a building with less post-earthquake residual deformation than typical non-centering systems.

Because the period shift is critical to the design and performance of the RSFJ, the analysis should be completed using a deformation-based

Making the Choice

Photo of a braced bay, showing the steel tie that minimizes the chevron to beam connectors. Photo courtesy of Aspect Structural Engineers.

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In the end, the friction damper and the glulam brace options were both eliminated early in the study. Glulam braces did not have a high enough response modification factor to suit the project’s high seismic demand, and given the preferred brace layout, this system would have driven up the cost of timber members, diaphragm detailing, and foundation design. Slip friction dampers were likewise eliminated once the perimeter moment frames were sized and found to be too costly and having too much architectural impact due to the larger members. With roughly comparable performances for both the Continued on page 32

minimize steel-to-timber connections at the top of each chevron, as shown previous page on the bottom left. A raised access floor above the CLT deck allowed this steel tie, as well as all steel collectors, to remain hidden.

While the design strategy was based on solid engineering principles and an understanding of timber mechanics, the Terminus team was then pleasantly surprised to learn that this approach was independently verified by research in New Zealand that was underway shortly before construction of Terminus began. A PhD candidate at the University of Canterbury, Wenchen Dong, tested the performance of BRB braces in mass timber frames with different connection approaches and published his PhD thesis in 2021. He found that the best connection performance came from tight fit pin connections, which demonstrated moment release and did not restrict frame rotation.

Getting it Built Braced frames in the finished building. Photo courtesy of Aspect Structural Engineers, copyright Dasha Armstrong.

Continued from page 30 BRB and the Tectonus RSFJ options, the decision came down to the cost of the systems. Ultimately, BRBs by Corebrace housed within a mass timber frame were selected as the building’s lateral system.

Design Approach to BRBs in a Timber Frame The design approach started with the BRB system detailing requirements included in the Canadian steel code (CSA-S16), many of which are similar to the requirements in AISC 341. Some of the code provisions are listed below: a) The BRB elements shall be able to resist, without buckling, the forces and deformations that develop in the brace at 2.0 times the seismic design story drift. b) The probable tensile and compressive resistances of the bracing members shall account for strain hardening (tension and compression) and friction (compression only) via adjustment factors. c) The resistance of brace connections shall equal or exceed the probable tensile and compressive resistances of bracing members. d) The resistance of beams, columns, and all connections other than brace connections shall equal or exceed the effect of gravity forces and the probable tensile and compressive resistances of bracing members. e) Columns in multi-story buildings shall be continuous over a minimum of 2 stories.

In addition to this code guidance, adding the following provisions to the timber detailing approach justified that the performance would be comparable to an all-steel system: f ) Ensure deformation compatibility can be achieved by minimizing rotational restraint at beam-to-column connections to avoid the timber frame behaving as a moment frame. Aligning work points, slotting steel elements, and using small-diameter connectors (see item “g”) all helped to achieve this critical rotational release. g) Maintain ductile timber connections via many small-diameter, tight-fit pins and tight-fit bolts, even though the timber connections were not designed to yield or dissipate energy. Doing so better allowed the team to establish the deformation compatibility identified above. h) Introduce a steel tie member at the braced frame beam to 32 STRUCTURE magazine

Most of the SFRS options considered above are not codified and needed to be submitted to the AHJ as a code alternate. Working closely with the City of Langford’s building department early in the design process allowed the team to identify and address their requirements for approving the system. In general, early engagement with the AHJ is ideal for introducing new concepts and code alternates and allowing them time to learn, ask questions, and provide the pathway for approval. On the Terminus project, the AHJ was satisfied that the detailing approach for the BRBs was adequate to allow the BRBs to be used with the full response modification factor without penalty.

The early work on design and analysis paid off with a very smooth construction period. To compress the schedule, the mass timber connection detailing was completed in parallel to the mass timber supplier (Structurlam) creating the 3D fabrication model and the shop drawings. Close coordination and regular communication with Structurlam were key to this synchronous approach.

The BRB braces and timber frames resulted in a beautiful, polished architectural finish with a high-performing, buildable, and efficient lateral system. This new hybrid BRB-timber system serves as a great example of thoughtfully combining engineering mechanics, knowledge of the materials, and aesthetics. Meanwhile, the investigation of the SFRS options available exposed the impressive range of possibilities that can be applied to mass timber construction.■ Full references are included in the online version of the article at STRUCTUREmag.org.

Project Team EOR: Mehrdad Jahangiri Project Manager: Ilana Danzig Design Engineer: Jackson Pelling Design Engineer: Brendan Fitzgerald

Ilana Danzig, P. E., S. E., P. Eng, Struct. Eng., M. Eng, is a structural engineer in Vancouver, BC, Canada and Principal at ASPECT Structural Engineers. With experience in all materials her specialty is mass timber design and detailing, and she works on mass timber projects throughout Canada and the United States both as Engineer of Record and Design Assist Engineer (ilana@aspectengineers.com).

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code UPDATES 2024 IBC Significant Structural Changes Roof assemblies (Chapter 15)—Part 1.

By John “Buddy” Showalter, P. E. and Sandra Hyde, P. E.

T

his multi-part series discusses significant structural changes to the 2024 International Building Code (IBC) by the International Code Council (ICC). Part 1 includes an overview of changes to Chapter 15 on roof assemblies. Only a portion of the chapter’s total number of code changes are discussed in this article. More information on the code changes can be found in the 2024 Significant Changes to the International Building Code, available from ICC (Figure 1).

Roof Drainage Scuppers Unless a roof is designed to support the load from water ponding on top of it, both the IBC and the International Plumbing Code (IPC) require that a roof drainage system be provided. This system must include primary and secondary drainage systems or scuppers to serve as the secondary (emergency overflow) drainage (Figure 2).

1502.3 Scuppers. Where scuppers are used for secondary (emergency overflow) roof drainage, the quantity, size, location and inlet elevation of the scuppers shall be sized to prevent the depth of ponding water from exceeding that for which the roof was designed as determined by Section 1611.1. Scuppers shall not have an opening dimension of less than 4 inches (102 mm).

The flow through the primary system shall not be considered when locating and sizing scuppers.

Change Significance: Because IBC Sections 1502.1 and 1502.2 require compliance with Chapter 11 of the IPC, the scupper requirements in IBC Section 1502.3 were considered redundant and unnecessary; therefore, this section was eliminated. Where a jurisdiction does not adopt the IPC, it is suggested that reference to the IPC be treated like any other referenced standard (IBC Section 102.4). IBC Section 101.4 states that provisions of the IPC would “be considered to be part of the requirements of this code to the prescribed extent of each such reference.” By following this path, the requirements specified in IPC Section 1106 for determining rainfall rates and storm drainage requirements are still the expected standard of compliance.

Wind Resistance of Aggregate-surfaced Roofs Parapets for aggregate-surfaced roofs provide wind resistance and reduce the probability of aggregate blow-off. IBC Table 1504.8 considers the aggregate size, roof height, wind speed, and exposure category to determine the minimum required parapet wall height for aggregate-surfaced roofs. The provisions are based on wind speeds for blowoff and only address smaller aggregate used for the surfacing of 36 STRUCTURE magazine

Figure 1 2024 Significant Changes to the IBC.

built-up roofs and sprayed polyurethane foam roofs, which are different systems than ballasted roofs.

1504.9 1504.8 Wind resistance of aggregate-surfaced roofs.

Parapets shall be provided for aggregate-surfaced roofs and shall comply with Table 1504.9 1504.8. Such parapets shall be provided on the perimeter of the roof at all exterior sides except where an adjacent wall extends above the roof to a height at least equivalent to that required for the parapet. For roofs with differing surface elevations due to slope or sections at different elevations, the minimum parapet height shall be determined based on each roof surface elevation, and at no point shall the parapet height be less than that required by Table 1504.8.

Exception: Ballasted single-ply roof coverings shall be designed and installed in accordance with Section 1504.5.

Figure 2 Scuppers used as secondary drains. Graphic courtesy of ICC.

TABLE 1504.9 1504.8 MINIMUM REQUIRED PARAPET HEIGHT (INCHES) FOR AGGREGATE-SURFACED ROOFS a, b, c, d, e AGGREGATE SIZE

MEAN ROOF HEIGHT (ft)

WIND EXPOSURE AND BASIC DESIGN WIND SPEED, V (MPH) Exposure Cd f Exposure B ≤ 95 100 105 110 115 120 130 140 150 ≤ 95 100 105 110 115 120 130 140 150

Body of table unchanged and not shown for brevity. a. Parapet height is measured vertically from the top surface of the coping down to the surface of the roof covering in the field of the roof adjacent to the parapet and outbound of any cant strip. b. a. Interpolation shall be permitted for wind speed, mean roof height and parapet height. Extrapolation is not permitted. c. b. Basic design wind speed, V, and wind exposure shall be determined in accordance with Section 1609. d. c. Where the minimum required parapet height is indicated to be 2 inches (51 mm), a gravel stop shall be permitted and shall extend not less than 2 inches (51 mm) from the roof surface and not less than the height of the aggregate. e. The tabulated values apply only to conditions where the topographic factor (Kzt) determined in accordance with Chapter 26 of ASCE 7 is 1.0 or where Kzt is incorporated in the basic wind speed in Section 1609. f. d. For Exposure D, add 8 inches (203 mm) to the parapet height required for Exposure C and the parapet height shall not be less than 12 inches (305 mm).

Change Significance: Revisions clarify how the protection of aggregate-surfaced roofs is to be applied. A parapet is not required on any side of a building’s roof where a wall extends to or above the height required for a parapet. This situation may occur when a building steps up with an additional story or stories above a lower section. An example would be where the roof of a two-story building segment steps up to a third story or more (Figure 3). Although not specifically stated, the intent is that the building wall is on the same building and not a

higher wall located on an adjacent building or neighboring property.

Note also that a structure can have multiple roofs at different elevations, and parapet heights are determined by evaluating each roof surface. All roofs on the building are not required to meet the most stringent condition. An example of this might be where the roof of a high-rise tower could face a substantially different wind impact than the roof of the lower base of the building. Therefore, each “roof ” is evaluated separately when applying IBC Table 1504.8.

The new exception to IBC Section 1504.8 points to the ballasted low-slope single-ply roof provisions of IBC Section 1504.5, where the ANSI/SPRI RP-4 Wind Design Guide for Ballasted Single-ply Roofing Systems standard addresses the edge height requirements. In this case, IBC Table 1504.8 is not applicable. The important distinction is that the size of the aggregate addressed by IBC Table 1504.8 and used as an aggregate surface for the roof is smaller (3/4 inch or less), while stone used as ballast per IBC Section 1504.5 is larger (11/4 to 2½ inches). This difference in aggregate size is established

by IBC Table 1504.8 reference to ASTM D1863 Specification for Mineral Aggregate Used on Built-up Roofs. Conversely, IBC Section 1504.5, and IBC Section 1507.12 which it references, use ASTM D448 Standard Classification for Sizes of Aggregate for Road and Bridge Construction or ASTM D7655 Standard Classification for Size of Aggregate Used as Ballast for Roof Membrane Systems to set aggregate size (Figure 4).

A new footnote “a” to IBC Table 1504.8 provides specific language to address how the parapet height will be measured. Specifically, the height is measured above the main surface of the roof and is not intended to be measured from the height of the cant strip.

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A new footnote “e” to IBC Table 1504.8 clarifies that the values in the table were established and intended to only apply to sites with no topographic effects. Therefore, the value of Kzt must be 1.0 or be incorporated in the basic wind speed determined in Section 1609. Where topographic effects exist at the site (e.g. Kzt > 1.0), it could result in increased wind speeds at the roof height and a more severe condition for aggregate blowoff than what was considered when the table was developed.

Reroofing

Figure 3 Parapet exempt at the adjacent wall.

scouring or blow-off of the aggregate surfacing, the impact of the wind flow over the parapet will affect the roof field without creating a wind problem right next to a parapet. While the cant strip is needed for drainage, it will not face the same wind impact when applying this blow-off requirement. However, code users need to realize that while IBC Section 1504.8 will not consider the height of the cant strip, IBC Section 705.12.1 sets the parapet height for fire resistance differently. Therefore, both sections must be considered before determining the required parapet height and from what point it is measured.

The revision to IBC Table 1504.8 footnote “b” (Footnote “a” in the 2021 IBC) clarifies that interpolation for the wind speed is permitted between any values listed in the table. The footnote also clarifies that users should not extrapolate values beyond the table limits. Additional research and evaluation are needed for any conditions beyond the limits of the table.

Aggregate-surfaced Roofs IBC Table 1504.8

ASTM D1863 #7 aggregate – 0.1875 to 0.5 in. #67 aggregate – 0.1875 to 0.75 in. #6 aggregate – 0.375 to 0.75 in.

1512.1 General. Materials and methods of application used for recovering or replacing an existing roof covering shall comply with the requirements of Chapter 15.

Exceptions: 1. Roof replacement or roof recover of existing low-slope roof coverings shall not be required to meet the minimum design slope requirement of ¼ unit vertical in 12 units horizontal (2-percent slope) in Section 1507 for roofs that provide positive roof drainage and meet the requirements of Section 1608.3 and Section 1611.2.

Ballasted Low-slope Roofs

IBC Section 1507.12.3

ASTM D448 ASTM D7655 #4 ballast – 1.25 in. #2 ballast – 2.5 in.

Figure 4 Aggregate size differences for ballasted versus aggregate-surfaced roofs.

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When a roof is replaced or recovered, it must comply with IBC Chapter 15 for the materials and methods used, with exceptions for low-sloped roofs and a waiver of the secondary drainage provision when reroofing an existing building where the roof drains properly. This section does not mandate that the entire roof be replaced but simply that the replaced portion comply with IBC Chapter 15. During the reroofing process, susceptible bays must now be analyzed for ponding instability (Figure 5).

IBC Section 1504.5

ANSI/SPRI RP-4 ASTM D7655 #4 ballast – 1.25 in. #2 ballast – 2.5 in. #4 paver ≥ 18 psf

Figure 5 Susceptible bays must be analyzed for ponding during reroofing.

2. Recovering or replacing an existing roof covering shall not be required to meet the requirement for secondary (emergency overflow) drains or scuppers in Section 1502.2 for roofs that provide for positive roof drainage and meet the requirements of Section 1608.3 and 1611.2. For the purposes of this exception, existing secondary drainage or scupper systems required in accordance with this code shall not be removed unless they are replaced by secondary drains or scuppers designed and installed in accordance with Section 1502.2.

is evaluated. With changes that have occurred in the code over the past few years (e.g., increased rainfall intensities and changes to drainage provisions), existing roof systems may not meet the current requirements. As an example of where IBC Section 1512.1 Exception 2 may be applied, if a roof does not have an adequate number of scuppers or they are inadequate in size to accommodate the generally required overflow requirements, then at that point, the roof would need to be evaluated for ponding instability, or additional secondary drainage capacity would need to be provided.

Change Significance: The exceptions in IBC Section 1512.1 were revised to emphasize that snow loads must be checked during the reroofing process and susceptible bays must be analyzed for ponding instability due to rain loads. The intent is to reduce the likelihood that inadequate drainage and ponding caused by new loading or roofing configurations might lead to either a roof collapse or failure. It is often difficult, if not impossible, to make changes to the roof slope or increase the drainage system when reroofing. While IBC Section 1512.1 Exception 1 previously would have allowed any existing roof that did not meet the 2-percent-slope provision to be reroofed, it now only allows reroofing to occur if the roof has both positive drainage and has been evaluated for ponding instability.

IBC Section 1512.1 Exception 2 recognizes that it could be possible that the existing secondary drainage or scupper system may not be capable of meeting the current code requirements. Therefore, this exception will still permit the reroofing project to proceed provided there is positive roof drainage, and the potential for ponding instability

Conclusion Structural engineers responsible for roof assembly design should be aware of significant structural changes in the 2024 IBC. The scupper requirements for secondary roof drainage have been removed from the IBC. New provisions include clarity regarding the wind resistance of aggregate-surfaced roofs where a parapet is not required on any side of a building’s roof where an adjacent wall extends to or above the height required for the parapet. Finally, susceptible bays must now be analyzed for ponding instability during the reroofing process.■ Look for more of the series in upcoming issues of STRUCTURE. John “Buddy” Showalter, P. E., (bshowalter@iccsafe.org) is Senior Staff Engineer and Sandra Hyde, P. E., (shyde@iccsafe.org) is Managing Director of ICC’s Consulting Group.

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iconic STRUCTURES San Francisco–Oakland Bay Bridge One of America's greatest bridges, carrying the heaviest traffic in Northern California.

By Roumen V. Mladjov, S. E., P. E.

T

he San Francisco–Oakland Bay Bridge (henceforth, Bay Bridge), completed November 12, 1936, during the Great Depression, remains one of the world’s greatest bridges. Together with its neighbor, the Golden Gate Bridge, it represents the culmination of more than 100 years of bridge engineering development in the US. After its opening, the bridge, which connects San Francisco and Oakland, soon became known as the “workhorse of Northern California,” carrying the heaviest traffic in the region. As part of Interstate 80, it remains the busiest traffic link in Northern California carrying on its two decks up to 280,000–300,000 vehicles per day, or more than 13,000 per peak hour.

The bridge consists of several structures with distinctly different systems, connected to form an 8.5 mile crossbay roadway, nearly 4.4 miles of it over water. The main portions of the original bridge were: • West Crossing Span: Twin suspension spans with two decks. Central spans of over 2,310 feet. Total length of nearly 2 miles from San Francisco to Yerba Figure 1 West Crossing of Bay Bridge (as seen from San Francisco).

Buena Island (YBI), see Figure 1.

• YBI Segment: A two-level tunnel and short concrete viaduct, with a length of 1,800 feet. with two consecutive suspension spans.

• East Crossing Span: Consisting of several different steel truss systems: The bridge is listed on the National Register of Historic Places four short 288-foot spans on YBI, followed by the 2,420-foot long (NRHP) with a note: “One of the largest and most important cantilever structure (Figure 3), then five deep through-truss spans historic bridges in the country.” From the graceful West Crossing of 509 feet each, fourteen deck-truss spans at 288 feet, and steel suspension structure, through the Yerba Buena Island tunnel and structures. More than 2 miles long crossing from YBI to Oakland. viaduct, to the steel cantilever truss section and truss spans, this entire The original bridge was designed by Ralph Modjeski and Charles bridge deserves to be remembered for these historic credentials. It Purcell, among others, was and it remains one and was built by the of the greatest engineerAmerican Bridge ing achievements of the Company from 1933 20th century. to 1936, for a total The engineers uticost of $77 million, lized state-of-the-art with steel from United techniques available in States Steel. Upon the 1930s. They used opening, it was the the highest-strength longest highway bridge construction steel availin the world. It had the able: nickel steel (55 second longest suspenksi) and silicon steel sion span (2,310 feet), (45 ksi) made up the the third longest cantilarger portion of the lever truss span (1,400 total steel employed feet), the deepest pier for the East Crossing. foundation below water Even the carbon steel (243 feet) and the largused in this bridge was est bored tunnel. The higher strength (37 West Crossing was also ksi) than that normally the only major bridge Figure 2 Proposed "Rush San Francisco Trans-Bay Suspension Bridge" (1913). used today. The West

40 STRUCTURE magazine

Figure 3 (Left) West Crossing Span as seen from above. With East Crossing in the distance beyond YBI. “Above San Francisco”, courtesy of Robert Cameron.

(Right) West and East Crossing Spans from above. Image courtesy of Google Earth.

Crossing suspension steel cables are 120 ksi. The structural steel used for the entire bridge amounted to 167,100 tons, or 125 psf.

Historical Development Already before the beginning of the 20th century, the need for a bridge was felt by Californians in the San Francisco Bay Area. The self-proclaimed Emperor Norton I, an eccentric resident of San Francisco, thrice decreed that a suspension bridge be constructed to connect Oakland with San Francisco, the last time on September 17, 1872.

Unlike most of Emperor Norton's eccentric ideas, his decree to build a bridge had wide public and political appeal. Yet, the task was a major engineering and economic challenge since the waters between San Francisco and Oakland were too wide and too deep to be tackled by the engineering technologies of the day. In 1921, over forty years after Norton's death, a tunnel tube was considered, but it became clear that it would be inadequate for vehicular traffic. Support for a trans-bay crossing finally grew in the 1920s (An

early idea for such a bridge is shown in Figure 2) with the increasing popularity and availability of the automobile. In April 1932, the preliminary plan and design of the bridge were approved, and preparation of the final design proceeded.

Original 1936 Bridge Construction began on July 9, 1933, after a groundbreaking ceremony attended by former president Herbert Hoover, who declared: “This marks the physical beginning of the greatest bridge yet erected by the human race.” The bridge opened after just over three years of construction, justifying President Hoover’s appraisal by instantly becoming one of the greatest bridges ever built.

Since the Bay Bridge spans a long distance, it necessitated specific bridge structures for its different portions. Intersecting the maritime route into the ports of both San Francisco and Oakland, it required a minimum clearance of 220 feet above water. The western section of the bridge, between San Francisco and Yerba Buena Island, presented

Figure 4 Cantilever bridge east of Yerba Buena Island (original Bay bridge).

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Figure 5 Original bridge traffic, 1936 to 1963.

an enormous engineering challenge. Here, the Bay was up to 100 feet deep in places and would require new foundation-laying techniques.

A single main suspension span 4,100 feet in length was considered but rejected. The solution was to construct a massive concrete anchorage halfway between San Francisco and Yerba Buena Island, and to build main suspension spans on each side of this central anchorage.

The West Span (original & current) is a continuous suspension structure with two suspension main spans of 2,310 feet, interconnected with a middle 2,490 foot span divided in two by the massive central anchorage. The total suspension part of the west span is 9,452 feet.

The continuous suspension bridge spans (Figure 3), are as follows: San Francisco end span

1,171 feet

First main span

2,310 feet

Middle double span with shared anchorage (51.82 m) at mid-span and total length of

2,490 feet

Third main span

2,310 feet

Yerba Buena Island end span

1,171 feet

Reaching the Yerba Buena Island, the road goes through a two-story Yerba Buena Tunnel, which is 76 feet wide, 58 feet high, and 540 feet long. It is the largest diameter transportation bore tunnel in the world.

The considerable amount of material excavated from boring the tunnel

Figure 7 Replacement East Span, 2016.

42 STRUCTURE magazine

Figure 6 Damaged East Span.

was used for part of the landfill north of Yerba Buena Island, creating the artificial Treasure Island.

The original East Span from Yerba Buena Island to Oakland with total length of 10,176 feet was a combination of a double cantilever truss structure (see Figure 4), five long-span through-trusses, and a multi-span steel truss causeway with spans of 492 feet. The cantilever truss span of 1,401 feet was the longest in the nation and the thirdlongest anywhere in the world.

The foundations of most of the original eastern section were on treated wood pilings. Because of the very thick layer of mud on the Bay bottom, it was not practical to reach bedrock, although the lower levels of the “mud” are quite firm. Long wooden pilings were crafted from entire old-growth Douglas Fir trees, which were driven through the soft mud to the firmer bottom layers. However, these foundations were later the main reason for the decision to replace the entire East Span.

Bridge Changes from 1936 to 1989 The original bridge initially carried automobile traffic on its upper deck, with trucks, tramways, buses, and commuter trains on the lower deck, (see Figure 5). In 1958, the lower deck was converted to all-road traffic as well. In October 1963, the traffic was reconfigured to one way traffic

Figure 8 Self-anchored suspension (SAS) bridge the dimensions are in meters!

on each deck, westbound on the upper deck, and eastbound on the lower deck, with trucks and buses now allowed on both decks.

East Span Replacement 2002–2016 Following the Loma-Prieta Earthquake 1989 On October 17, 1989, a 6.9M earthquake hit Northern California. It had an epicenter on the San Andreas Fault, roughly 56 miles south of San Francisco, and lasted 20 seconds. As a result, a small 50-foot long upper deck portion of the East Span slipped from one of its supports and dropped onto the lower deck (Fig.6), necessitating month-long repair and initiating a heated discussion for the retrofitting or replacement of the bridge. The structural decision was to replace the entire East Span of the bridge completely, mostly because of the inadequate support capacity of its timber pile foundations and the very difficult retrofitting required for such foundations.

After several years of discussion, planning and design, construction on the new east crossing finally began in January 2002, and was completed after 14 years of construction in October 2016, for $6.5 billion for a length of nearly 2.2 miles, but without adding a single additional traffic lane to relieve the heavy congestion on the bridge. Partial traffic was allowed starting in 2013.

The East Span replacement was designed by T.Y. Lin International, Moffat & Nichol Engineers, Weidlinger Associates, and Donald MacDonald Architects. It is comprised of a single-tower, self-anchored suspension (SAS) steel span of 1266 feet, see Figures 7 and 8, and a 14-span (525 feet each) concrete skyway, see Figure 9. The new crossing has added shoulders and a bicycle lane, but there is no bicycle lane on the West Span.

Unlike the double-deck western and the original eastern section of the bridge, the new replacement East Span (eastern part) consists of two single deck bridges carrying both eastbound and westbound lanes. Following completion of the replacement structure, demolition and removal of the old East Span began in 2016 and was completed on September 8, 2018, at an estimated cost of at least $250 million.

Replacement East Span/Technical data: Cable-stayed “Signature” bridge with a main span of 1,263 ft a back span of 590 ft; and a total length of 2,048 ft.

Skyway, was constructed with multiple 525 ft spans of comprised of precast post-tensioned concrete segments. The total concrete utilized is 450,000 cubic yards. The total weight of the Skyway structure is 8.48 t/m2.

Bridge Element

Original East

New East Bridge

Total Length (meters)

3,377

3,513

Main Span (meters)

427

385

Secondary Spans (meters)

155

180

Traffic Lanes

10

10

Vehicles per Day

280,000

320,000 (Aug. 2019); 305, 300 (Apr. 2023)

Construction Time (years)

3.5

14

Completed/Estimated

1936

2016

Steel (kg/m2)

416

1,843

Cost US

$ 78 M (non-comparable)

6,450 M

Seismic Safety Improvements to the West Span The West Span (and its western approach) underwent seismic improvements in a five-year project beginning in 1999, at a reported cost of approximately $759 million. The improvements included massive rollers installed between the roadway and bridge supports and 96 new viscous dampers inserted at critical points to damp movement. The bridge’s twin suspension spans were strengthened by adding new steel plates and replacing N OVE M B ER 2023

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Figure 9 East Span, Skyway under construction 2002–2016.

Figure 10 The illuminated West Span as seen from the Embarcadero Center, San Francisco, with LED lights complementing the pre-existing permanent lights.

half a million original rivets with almost twice than many modern highstrength bolts. New diaphragm bracing was added under both decks, and all the “laced” truss diagonals connecting the upper and lower road decks were replaced. In total, the project added about 8,500 tons.

Traffic The San Francisco-Oakland Bay Bridge is the region's “workhorse bridge,” carrying more than a third of the traffic of all the state-owned bridges combined. By 1990, Bay Bridge traffic had already reached 280,000 vehicles per day, which later increased to 320,000, creating one of the worst traffic congestions in the nation. Traffic is predicted to grow further to 388,000 by 2040. During the Covid epidemic and 44 STRUCTURE magazine

“remote working from home” there was some reduction of the heavy traffic, however already in 2023 there were signs of traffic increasing to near pre-Covid numbers. In April 2023 the traffic on the bridge is already 96% of the maximum pre-covid from early 2020.

Nearly 90 years after its completion, despite challenges and modifications, the Bay Bridge remains both an emblematic landmark for the San Francisco Bay Area and a practical and indispensable feature of its daily life. It is also a milestone in the achievement of the Art of American Engineering during the 20th century.■ Roumen Mladjov, S. E., P. E., Roumen’s main interests are structural and bridge development, structural performance, seismic resistance, efficiency, and economy. (rmladjov@gmail.com).

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historical STRUCTURES 19th Century Mississippi River Bridges #9 Winona bridge, Minnesota/Wisconsin 1871.

By Frank Griggs, Jr., Dist. M. ASCE, D. Eng,. P. E., P. L. S.

T

he Winona and St. Peter Railroad (originally the Transit Railroad Company) was founded in 1862 and planned to run westerly from Winona on the Mississippi River to St. Peter, about 125 miles away, reaching that point in 1871. From there, it reached Watertown, South Dakota, in 1873. In 1867, it was purchased by the Chicago and North Western Railway that was chartered on June 7, 1859, and from Chicago, it reached the Mississippi at a point in Wisconsin east of Winona, Minnesota, in 1873. It continued its route north to St. Paul, connecting with the Northern Pacific Railroad heading for the West Coast at Puget Sound. However, a bridge connecting the two lines, now under the control of the Chicago and North Western Railway, was needed. After a five-year battle, approval was given to build the bridge in accordance with the 1866 federal law prescribing the requirements for a low-level and a high-level bridge as described for the Quincy and other bridges. Section 8 of the July 25, 1866 Act stated,

And be it further enacted That the Winona and Saint Peter Railroad Company, a corporation existing under the laws of the state of Minnesota, is hereby authorized to construct and operate a railroad bridge across the Mississippi River between Rebuilt 363’ long swing span in iron (from an advertisement). the city of Winona in the state of Minnesota and the opposite bank of the said river, in the state of Wisconsin, with the consent allowing for 160’ clearance between the swing pier and rest piers. Since it was of the legislature of the states of Minnesota and Wisconsin and said for a low-level bridge, the 1866 Law required a clearance of thirty feet above bridge by this section authorized is hereby declared a post route, and the low water mark and not less than ten feet above the high water mark. subject to all the terms, restriction and requirement contained in the The Winona and St. Peter Railway then ran along the westerly bank of the foregoing sections of this act. river just above the water line heading north to Minneapolis. This required With the approval of Minnesota and Wisconsin, they began the preliminary a long fill and curved trestle to raise the tracks to the bridge level. Once at planning for the bridge. They chose a low-level bridge with a swing bridge, the required level, there was a short single 80’ span pony truss followed by a 360’ wooden swing span on the Post pattern. This was followed by two Howe 250’ through trusses with curved top chords. The bridge was 980 feet long across the main channel of the river. This was followed by a trestle across Latsch Island and a three-span Howe through truss over the westerly branch of the Mississippi to Wisconsin. This was followed by a trestle on a curve and a straight line to another Howe through a truss over a slough. The American Bridge Company did not build the portion from the island to touch down in Wisconsin.

The main structure on the bridge was a long wooden swing span near the Winona bank of the river. Since the American Bridge Company from Chicago had adopted the Post Truss for its standard bridge, they chose that pattern for this span with Howe deck and through trusses for the shorter fixed spans. To cut down on costs, they designed the span in wood and iron with the posts inclined towards the center instead of vertically and used a curved top chord in wood.

The bridge was finished on May 26, 1871, and the Winona Republican newspaper wrote, “The bridge is both imposing and symmetrical in its appearance, and it possesses all the strength Minnesota end of the bridge with the swing span and three-span bridge over the easterly and solidity that the most scientific combination of stone, iron, branch of the river. Note the swinging curve old wooden bridge shown on the left to the island.

46 STRUCTURE magazine

and wood, can impart to any structure of the kind.” But on June 1, the St. Cloud Democrat wrote, “The first train of cars crossed the Mississippi last Thursday evening [May 25, 1871]. A grand celebration of the completing of the bridge was to have been today [June 1, 1871], but last Saturday, the eastern half of the draw gave way with a train of six flat cars loaded with stone…The length of the draw was 360 feet, and the portion which gave way was on the Wisconsin side about 180 feet long. The bridge had not been fully adjusted and had not been accepted by the railroad company.”

Another paper wrote in part, “One must be horrified when he contemplates what would have been the awful spectacle had been away with the passenger train scheduled few hours before instead of the freight train as it did. If the company from Chicago which built the Winona Bridge was guilty of deception or careless work, they ought to smoke for it and be driven out of their homes. We are cursed with work of this kind.” Another paper wrote, “It was the frailest thing, the smallest piers and cheapest structure on the river, costing only $130,000 while the magnificent Dubuque bridge cost $600,000.” It was discovered that the bridge builder was not Entire length of the bridge from Winona on top of the image to low lowlands in Wisconsin. fully responsible for the failure, and the bridge tender, new to his job, had not properly closed the bridge, and the train derailed, causing the out of business in 1878 after a failure of their attempt to build a bridge across failure. The American Bridge Company rebuilt the swing span in iron, and the Hudson River at Poughkeepsie, New York. it was swung on January 13, 1872. On January 21, 1872, the bridge was The swing span was replaced in 1899, and the rest of the bridge was replaced reopened for traffic. over time with a series of plate girder spans over much of its length. The bridge The Bridge Company had an advertisement prepared with the image above was abandoned in 1977. A portion of the plate girder spans remain in place.■ that gave the following data on their bridge, “Swing span, Post’s Patent Dr. Frank Griggs, Jr. specializes in the restoration of historic bridges, having Diagonal Iron Truss, with [steam] engine, 360 feet; Fixed spans Howe Truss restored many 19th Century cast and wrought iron bridges. He is now an Arched Top chords 250 feet each; Fixed Howe Truss 80 feet. Total length of Independent Consulting Engineer (fgriggsjr@twc.com). bridge over main channel, 940 feet.” The American Bridge Company went ADVERTISEMENT–For Advertiser Information, visit STRUCTUREmag.org

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structural CONNECTIONS Vibration of Steel Joists with Flush Frame End Connections

Steel joists with flush framed end connections enhance floor vibration properties.

By David Samuelson, P. E., Brad Davis, Ph. D., S. E., and Thomas M. Murray, Ph. D., P. E.

R

ecent floor vibration measurements and research completed by Vulcraft have shown that flush frame joist end connections can have a significant positive impact on a floor’s vibrational response when subjected to walking excitation. This article shows how the frequency and effective mass of the floor bay can Figure 1 Joist Seat and Flush Frame Joist-to-Girder Connections. be changed dramatically when switching from traditional bearing seats to flush frame connections AISC DG11 Provisions for Flush developed by Vulcraft. Manual calculation procedures found in the American Institute of Steel Construction (AISC) Design Guide Frame Joist Connections 11, 2nd Edition, Vibrations of Steel-Framed Structural Systems Due to Human Activity and the Steel Joist Institute’s (SJI) Technical Following is an overview of how the applicable provisions for Digest 5, Vibration of Steel Joist – Concrete Slab Floors (hereafter walking excitations contained within AISC DG11 apply to flush referred to AISC DG11 and SJI TD5) currently do not address frame joist connections. Equations are referenced to AISC DG11. steel joists with flush end connections. This article demonstrates how the provisions in AISC DG11 and SJI TD5 can be used for Tolerance Criterion flush frame connections.

With traditional joist-bearing seats, the steel deck-supported concrete The recommended criterion for low-frequency building floors is as slab is above the top of the girder by the height of the seat, as shown follows: the floor system is satisfactory if the peak acceleration, ap, in Figure 1(a). With flush frame joist connections, the slab is directly due to walking excitation as a fraction of the acceleration of gravity, in contact with the girder, as shown in Figure 1(b). g, determined from Figure 2 shows flush frame joist connections to a girder and a joist ap Po e -0.35f girder. With this type of connection, a vertical plate at the end of (AISC DG11, 2nd Ed., Eqn. 4-1) g = bW the joist is field bolted to a simple shear connection at the support.

For a wide flange-to-girder connection, the concrete slab is also does not exceed the tolerance acceleration limit, αo/g [0.005 (0.5%g) directly attached to the top flange of the girder, as shown in for offices or residences], where Figure 3. For these connections, whether the slab is connected P0 = amplitude of a 65-pound (lb.) driving force using shear studs, ShearFlex® screws, or spot welds, the girder fn = fundamental natural frequency of a beam or joist panel, acts fully composite when subjected to walking excitation. AISC a girder panel, or a combined panel, as applicable, DG11 has guidance for predicting the natural frequency and Hertz (Hz) effective weight associated with floor acceleration due to walkβ = damping ratio ing and rhythmic activities for these connections. When steel W = effective weight supported by the beam or joist panel, girder joists are fabricated with flush frame connections, the vibration panel, or combined panel, as applicable, lb. behavior is similar to hot-rolled beams supported by hot-rolled or built-up girders.

Frequency The intent of this article is to 1) show the advantage of flush frame joist connections for controlling floor vibration caused by The girder or joist girder natural frequency is estimated using the occupant activities and 2) provide the structural engineer with fundamental natural frequency equation of a simply supported beam guidance for calculating the girder or joist girder natural frequency, with a uniform mass: fg, and effective joist panel weights, Wj, for flush frame-to-girder/ g fn = 0.18 T joist girder connections using the AISC DG11 provisions.

(AISC DG11, 2nd Ed., Eqn. 3-3) n

48 STRUCTURE magazine

inertia somewhat to significantly less than the full composite moment of inertia of the girder or joist girder.

The effective moment of inertia of joist girders supporting traditional joist seats is estimated using Ig = Cr Ichords +

Ie - Cr Ichords (AISC DG11, 2nd Ed., Eqn. 3-10) 4

where Cr = coefficient from AISC DG11, 2nd Ed., Eqn. 3-9a Ichords = moment of inertia of the chord areas, in.4 Ie = effective composite moment of inertia from AISC DG11, 2nd Ed., Eqn. (3-7), in.4

Ie =

c Ichords

1 + I1

(AISC DG11, 2nd Ed., Eqn 3-7)

comp

where Ichords = moment of inertia of the chord areas alone, in.4 Icomp = fully composite transformed moment of inertia of the slab and chord areas, in.4 Similarly, the effective moment of inertia of hot-rolled or built-up girders supporting traditional joist seats is estimated using I g = Ix +

Icomp - Ix 4

(AISC DG11, 2nd Ed., Eqn. 3-11)

where Ix = moment of inertia of the girder, in.4 Icomp = fully composite transformed moment of inertia of the slab and girder areas, in.4

Effective Weight

Figure 2 Flush Frame Joist End Connections.

where fn = fundamental natural frequency, Hz g = acceleration of gravity, 386 inch/second.2 ∆ = mid-span deflection of the member relative to its supports due to the supported weight: that is, ∆ = 5wL4/(384EsIt), inches (in.). where Es = modulus of elasticity of steel, 29,000 kips per square inch (ksi) It = transformed moment of inertia; effective transformed moment of inertia if shear deformation is included; reduced transformed moment of inertia to account for joist seat flexibility, in.4 w = uniformly distributed weight per unit length (actual, not design, dead and live loads) supported by the member, kip/in.

L = member span, in.

Human-induced loads typically cause mid-bay displacement amplitudes smaller than 0.01 in., implying very low horizontal shear force amplitudes (Figure 4) between the steel framing members and the slab. When the deck is in direct contact with the top of the member, deck fasteners, including spot welds and screws, provide enough slip resistance to ensure fully composite behavior. Members with physical separations between the member and the slab, e.g., a girder or joist girder supporting open-web steel joists with traditional seats, behave as partially composite members with an effective moment of

The effective panel weight is estimated by determining the effective panel weights for the beam or joist panel (Wj) and girder panel (Wg) modes separately and then combining them in proportion to their flexibilities.

Tg Tj W = T + T W j + T + T Wg (AISC DG11, 2nd Ed., Eqn. 4-5) j

g

j

g

Where ∆j and ∆g = mid-span deflections of the beam or joist and girder, respectively, due to the weight supported by the member, in.

Wj and Wg = effective panel weights from AISC DG11, 2nd Ed., Eqn. (4-2) for the beam or joist and girder panels, respectively, lb.

Figure 3 Single Plate Beam-to-Composite Girder Web Connection.

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The effective panel weights for the beam or joist and girder panel modes are estimated from W = wBL

(AISC DG11, 2nd Ed., Eqn. 4-2)

where w = supported weight per unit area, pounds per square foot (psf ) L = member span, feet (ft.)

B = effective panel width, ft.

For the beam or joist panel mode, the effective width is Bj = Cj(Ds/Dj)1/4 Lj ≤ 2/3 × Floor Width (AISC DG11, 2nd Ed., Eqn. 4-3)

where Cj = 2.0 for joists or beams in most areas = 1.0 for joists or beams parallel to a free edge (edge of balcony, mezzanine, or building edge if cladding is not connected) Ds = slab transformed moment of inertia per unit width, in.4/ft 3 = approximately 12d e / ^ 12 n h in.4/ft, or from a deck manufacturer’s catalog de = effective depth of the concrete slab, taken as the depth of the concrete above the deck plus one-half the depth of the deck, in. n = dynamic modular ratio = Es/(1.35Ec) Es = modulus of elasticity of steel, 29,000 ksi Ec = modulus of elasticity of concrete, ksi Dj = joist or beam transformed moment of inertia per unit width, Ij/S, in.4/ft Ij = transformed or effective moment of inertia of the beam or joist, in.4 S = joist or beam spacing, ft Lj = joist or beam span, ft Floor width is the distance perpendicular to the span of the beams or joists in the bay under consideration over which the structural framing (beam or joist and girder size, spacing, length, etc.) is identical or nearly identical in adjacent bays.

For the girder panel mode, the effective width, except for edge girders, is 1/4

nd

Bg = Cg(Dj/Dg) Lg ≤ 2/3 × Floor Length (AISC DG11, 2 Ed., Eqn. 4-4)

Figure 4 Horizontal Shear Force at Deck-to-Girder Flange.

50 STRUCTURE magazine

where Cg = 1.6 for girders supporting joists connected to the girder flange with traditional joist seats = 1.8 for girders supporting joists or beams connected to the girder web Dg = girder transformed moment of inertia per unit width, in.4/ft = Ig divided by the average span of the supported beams or joists, in.4/ft Ig = effective transformed moment of inertia of the girder Lg = girder span, ft Floor length is the distance perpendicular to the span of the girders in the bay under consideration over which the structural framing (beam or joist and girder size, spacing, length, etc.) is identical or nearly identical in adjacent bays.

Where beams or joists are continuous at their supports and an adjacent span is greater than 0.7 times the span under consideration, the effective panel weight, Wj, can be increased by 50%.

This liberalization also applies to rolled sections shear-connected to girder webs but not to joists connected only at their top chord with traditional bearing seats.

Beam to Girder Web Connection The beam-to-girder web connection, shown in Figure 3, is a basic connection in AISC DG11. For this connection: • The girder is considered to be fully composite because the deck-togirder top flange connection at the steel deck-to-top beam flange interface provides adequate resistance to the vibration-generated horizontal shear force shown in Figure 4. The connection can be spot welds, Vulcraft’s Ecospan ShearFlex® screws, or standard welded steel shear studs.

• Cg = 1.8 when determining the girder panel mode effective width using AISC DG11 Eqn. 4-4.

• If an adjacent beam or joist span is greater than 0.7 times the span, then the effective joist panel weight, Wj, from AISC DG11 Eqn. (4-2) can be increased by 50%.

Traditional Bearing Seat Connection of Steel Joists to a Girder or Joist Girder The traditional bearing seat on a steel joist is fabricated utilizing two clip angles attached to the double-angle joist top chord, as shown in Figures 1(a) and 5(a). The traditional bearing seat provides less horizontal shear continuity than if the slab was directly connected to the girder top flange, as shown in Figure 5, so the fully composite transformed girder moment of inertia does not apply. Also, because the traditional joist-bearing seat does not transfer joist end moments across the girder, there is assumed to be no participation of the mass of adjacent bays in resisting walker-induced vibration.

Thus, for this connection: 1. The separation of the concrete slab and the girder top flange or joist girder top chord results in partial composite action, and the moment of inertia of the girder or joist girder supporting joist seats is determined using AISC DG11 Eqn. 3.11 or Eqn. 3.10, respectively. The primary result is a reduction in the natural frequency of the girder mode.

Vulcraft Sponsored Testing

Figure 5 Traditional Joist Bearing Seat Connection Deformation.

2. The reduced connection stiffness requires that the coefficient Cg = 1.8 in AISC DG11 Eqn. 4-4 be reduced to 1.6 when traditional joist seats are used, resulting in a slight reduction in the girder mode effective weight, Wg. 3. The non-participation of mass in adjacent bays means that an increase in effective joist panel weight, Wj, is not considered; that is, the 50% increase in joist panel weight, as recommended for shear-connected beam-togirder connections in AISC DG11 Section 4.1.2, is not applicable.

Flush Frame Joist Connection-to-Girder or Joist Girder The flush frame joist-to-girder and joist-girder connections shown in Figure 2 are equivalent to the typical beam-to-girder web connection shown in Figure 3.

For these connections, the following apply and are illustrated in Figure 6.

• A hot-rolled or built-up girder is considered to be fully composite.

A joist girder has the effective moment of inertia computed using AISC DG11 Eqn. 3-7.

• Cg = 1.8 when determining the girder panel mode effective width using AISC DG11 Eqn. 4-4.

• If an adjacent joist span is greater than 0.7 times the joist span, the effective joist panel weight Wj from AISC DG11 Eqn. 4-2 can be increased by 50%.

Wj = 1.5wjBjLj

Vulcraft sponsored a testing program to investigate the vibration performance of a composite joist floor system with flush frame composite joist connections. At the time of the testing, the building shown in Figure 7 was under construction. The tested floors supported no superimposed mass and were in the bare slab condition, as shown in the photograph. The purpose of the testing was to experimentally determine the dynamic properties of the floors and vibration response due to walking for comparison with finite element analysis and AISC DG11 manual calculation prediction methods.

The finite element analysis method accurately predicted the responsive natural frequencies and mode shapes in the tested bays. On average, the predicted natural frequencies were 5% higher than measured. The AISC DG11 manual method predicted natural frequencies were, on average, 10% higher than measured.

Accelerations due to walking were measured in three bays. Even though the floor was in the bare slab condition without nonstructural components that add damping and mass, the vast majority of accelerations were far below the recommended tolerance limit for quiet office spaces.

The tested bays were highly resistant to walking-induced vibrations and performed well in service. This Vulcraft-sponsored testing program showed that the acceleration prediction methods used in AISC DG11 could be confidently used for vibration analysis of composite joists with flush frame joist connections.

Example: Standard Joist Seats vs. Joist Flush Frame Connections and Hot-Rolled Beam Framing Connections The primary objective of this example is to show the effect on floor vibration response when flush frame joist connections are used for LH-series and CJ-series joists. A second objective is to show how a floor with flush frame connected joists compares to one with hot rolled members.

The exterior bay framing in Figure 8 is evaluated for vibrations due to walking in a modern (electronic) office. The building has joist spans of 45 feet-32 feet-45 feet and girder spans of 30 feet. The joist

Figure 6 Special Considerations for Flush Frame Connections.

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Figure 7 View of Tested Floor Having Flush Frame Joist Connections.

spacing is 72 inches. There are cold-formed steel walls above and below the spandrel girders. The exterior cladding is connected to the girders as well. Therefore, the spandrel girders are considered to be equivalent to exterior walls (infinitely stiff in the vertical direction) for vibration analysis. The floor slab has a total concrete depth of 5-1/2 inches, 2-inch deep composite deck, and normal weight concrete with f ʹc = 3.50 ksi.

The 45 ft bays have a superimposed dead load of 15 pounds per square foot (psf ) and a design live load of 50 psf (reduced to 45 psf for the secondary members). The resulting design dead plus live load, not including the weight of the secondary member itself or any bridging, is 696 pounds per linear foot (plf ). The loads for the 32-foot bay are 15 psf superimposed dead load and 100 psf live load.

Figure 8 Partial Framing Plan.

52 STRUCTURE magazine

The secondary joist and hot-rolled members shown in Table 1 satisfy strength and live load deflection requirements. The secondary joists with bearing seat connections are 28LH724/270 (required ASD 724 plf total load and 270 plf live load) joists and 28CJ716/270 joists. The secondary joists with flush frame connections are 28LH725/270 and 28CJ717/270. Note, from Table 1, the total loads shown in the joist designation include the weight of the joist. The non-composite interior girders are W27×84 sections, and the composite girders are W24×55 sections. For comparison with hot-rolled framing, the bay was analyzed for W18×35 composite beams and W24×55 composite girders.

Vibration response analyses of the bay with LHseries and CJ-series joists with standard seats, and flush frame connections, and hot-rolled framing, were conducted. For the vibration analyses, the live load is 8.0 psf, and the superimposed dead load is 4.0 psf (ceiling plus ductwork per DG11 Section 3.3). The secondary member and girder effective moments of inertia, Ieff , are per the “AISC DG11 Provisions” section above. The assumed damping ratio is 0.025, as recommended in AISC DG11 for electric office fit-out with ceiling and ductwork below. The results of the five analyses are summarized in Table 1. (The software FloorVibe v3.1, floorvibe.com, was used to generate the analysis results.)

28LH Joists and W27×84 Non-Composite Framing, Standard Bearing Seats vs.

Flush Frame Connections The predicted frequencies and acceleration ratio for the bay framing with 28LH724/270 and 5-inch deep standard joist seats are joist frequency 4.20 Hz; girder frequency 5.98 Hz; bay frequency 3.44 Hz, and acceleration ratio, ap/g times 100% = 0.606% (0.606%g). The recommended tolerance limit in AISC DG11 is 0.50%g which is less than the predicted acceleration ratio and therefore is not an acceptable design.

For the same framing, but with flush frame connections, the frequencies are 4.20 Hz, 7.63 Hz, and 3.68 Hz, and the acceleration ratio of 0.387%g, which is less than the AISC DG11 limit and is an acceptable design.

The reduction in the predicted acceleration with the use of flush frame connections is because of (1) an increase in the effective girder moment of inertia, Ieff ,g ,( 5183 in.4 to 8434 in.4), which in turn results in an increase in bay frequency (3.44 Hz to 3.68 Hz), and (2) an increase in the Continuity Factor from 1.0 to Continued on page 54

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LH-Series Joists

CJ-Series Joists

Hot-Rolled Sections

Secondary Member

28LH724/270 Non-Composite 27 plf

28LH725/270 Non-Composite 28 plf

28CJ716/270 Composite 19 plf

28CJ717/270 Composite 20 plf

W18×35 Composite 35 plf

Girder

W27×84 Non-Composite

W27×84 Non-Composite

W27×84 Non-Composite

W24×55 Composite

W24×55 Composite

Standard Seats1

Flush Frame

Standard Seats1

Flush Frame

Shear Framing

fj, Hz

4.20

4.20

3.69

3.69

4.03

fg, Hz

5.98

7.63

6.04

5.76

5.65

Connection to Girder

2

2

2

2

fn, Hz

3.44

3.68

3.15

3.11

3.28

ap/g × 100%

0.606 >0.500 NG

0.387 <0.500 OK

0.641 >0.500 NG

0.464 <0.500 OK

0.450 <0.500 OK

Ieff,j, in.4

1961

1964

1477

1479

1840

Ieff,g, in.

5183

8434

5184

4669

4670

Cont. Factor

1.0

1.5

1.0

1.5

1.5

Cg

1.6

1.8

1.6

1.8

1.8

Wj, Kips

134

200

140

211

208

Wg, Kips

115

115

113

112

116

W, Kips

129

185

135

189

183

4

5 in. standard joist-bearing seats ASD Total Load/Live Load. Total Load = 696 plf + joist weight plf + bridging weight plf.

1 2

Table 1 Comparison of Designs.

Continued from page 52 1.5, which results in a significant increase of the effective weight, (129 kips to 185 kips), as shown in Table 1.

Note: If the girders supporting the flush frame connected joists are composite, the predicted acceleration ratio with W24x55 composite girders is 0.452% - a satisfactory result.

28CJ Joists, Standard Seats vs.

Flush Frame Connections As shown in Table 1, the predicted acceleration decreases from 0.641%g to 0.464%g < 0.500%g when 28CJ joist end connections are changed from 5 in. standard joist seats to flush frame connections. The causes of the reduction are the same as explained for the 28LH joists and W27×84 non-composite framing – see Table 1 for related values.

Hot Rolled Framing The predicted frequencies and acceleration ratio for the bay framing hot-rolled beams and girders are a beam frequency of 4.03 Hz; a girder frequency of 5.65 Hz; a bay frequency of 3.28 Hz, an acceleration of 0.450%g < 0.500%g. It is noted that the secondary member weight, 35 plf, is significantly larger than for the joist secondary members.

Summary AISC DG11 and SJI TD5 provide guidance for predicting the natural frequency and effective panel weight for traditional steel joist seats supported by 54 STRUCTURE magazine

hot-rolled and built-up girders or joist girders. This article provides recommendations for determining the floor's natural frequency and effective panel weight when steel joists with flush frame joist connections are utilized.

With flush frame steel joist connections, the steel deck-supported concrete slab is in direct contact with the top of the girder. This results in the girder acting fully composite, as with hot-rolled beams framed into the web of a hot-rolled or built-up girder. Given the rotational restraint at the girder provided by the steel joist flush frame connections, vibrational energy is transferred across the girder over into the adjacent bay, increasing the slab effective weight by 50%, which in turn reduces the predicted acceleration ratio, ap/g, a significant benefit of the flush frame joist connection.

Utilizing flush frame joist connections, as opposed to increasing the concrete slab mass or steel joist depth to meet floor vibration requirements, can reduce the overall cost of the building. In areas where seismic forces can govern a building design, decreasing the concrete floor slab weights can reduce the seismic loads to that floor along with reducing the resulting foundation loads.■ Full references are included in the online version of the article at STRUCTUREmag.org.

David Samuelson, P. E., Structural Research Engineer, Vulcraft- Verco (Retired, Lake Andes, SD) Brad Davis, Ph. D., S. E., Associate Professor of Civil Engineering, Department of Civil Engineering, University of Kentucky, Lexington, KY. Thomas M. Murray, Ph. D., P. E., Emeritus Professor of Structural Steel Design, Department of Civil and Environmental Engineering, Virginia Tech, Blacksburg, VA.

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Rich responsive graphics. Simple intuitive inputs.

Detailed calculations. Project based organization. 41 different component type designers. Current/historical Section libraries. Section/Geometry optimization.

Applies preset and/or custom load combinations. FEM driven design. Powerful Revit integration. Simple, easy to use, reasonably priced.

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Steel Tube Institute Phone: 847-461-1701 Email: hssinfo@steeltubeinstitute.org Web: steeltubeinstitute.org Product: HSS Connex Description: The Steel Tube Institute’s HSS Connex is a free web-based tool that offers designers a simple way to check HSS wall thicknesses to aid in efficient HSS connection design. These local limit state checks are helpful to ensure HSS wall thicknesses are adequate at connections.

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Product: Tekla Structural Designer Description: With Tekla Structural Designer, engineers have the power to analyze and design multi-material buildings efficiently and cost effectively. Fully automated and packed with unique features for optimized concrete and steel design, Tekla Structural Designer helps engineering businesses win more projects and maximize profits.

Product: Tekla Tedds Description: Tekla Tedds automates repetitive and error prone structural and civil calculations, allowing engineers to perform 2D frame analysis, access a large range of automated structural and civil calculations to US codes and speed up daily structural calculations.

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by EORs to automate the creation of early connected models. Key features include detecting and resolving issues early and rapidly responding to design changes. To learn more about latest capabilities you can watch this recent webinar on our YouTube channel: https://www.youtube.com/ watch?v=HSRkWzzEnKM

RedBuilt Phone: 866-859-6757 Email: info@redbuilt.com Web: https://www.redbuilt.com/ Product: RedSpec™ Description: A convenient, user-friendly design program that lets you quickly and efficiently create floor and roof design specifications using Red-I™ Joists, RedBuilt™ Open-Web Trusses, RedLam™ LVL, Glulam Beams, and dimensional lumber. RedSpec™ is provided free of charge to registered users. For support, contact us at RedSpec@RedBuilt.com

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STRUCTUREmag.org STRUCTURE magazine

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SEICon24 at NASCC: The Steel Conference Technical Program at a Glance San Antonio, Texas | March 20 –22, 2024 | www.seicon24.org

KEY:

BLAST

BRIDGES

BUILDINGS

BUSINESS/ PROFESSIONAL PRACTICE

CAREER/LEADERSHIP

CLIMATE CHANGE/ SUSTAINABILITY/ LIFE CYCLE ASSESSMENT

CODES AND STANDARDS

WEDNESDAY, MARCH 20, 2024 10:15–11:15 AM

11:30 AM –12:30 PM

Designing Facades for Damage - Delegated Design for Hurricane & Blast Resistance

Fastening Forward: Recent changes, current gaps, and ongoing challenges in concrete connections

The New Disproportionate Collapse Mitigation Standard

Understanding How You’re Cost Accounting and Financial Systems Change Your Organizations Behavior

Development of Design Criteria for Functional Recovery: 2026 NEHRP and ASCE 7-28

Reconnaissance Observations from the Kahramanmaras Turkey Earthquake Sequence

THURSDAY, MARCH 21, 2024 9:45–10:45 AM

11:00 –12:00 PM

The Future of Structural Engineering & Embodied Carbon Workshop - Part 1

The Future of Structural Engineering & Embodied Carbon Workshop - Part 2

Updates to ASCE 7 Chapter 5 Flood Loads

Education and CROSS - “A Win-Win”

Structural Reliability in Practice

Loads on Temporary Structures - 2024 IBC and ASCE 7-28

FRIDAY, MARCH 22, 2024 8:00–9:00 AM

9:15–10:15 AM

From New to Normal: Bringing Ideas to Life

New Build or Retrofit - educating (and education of) the structural engineering profession about JEDI

Stainless Steel Design with ASCE 8-22

Machine Learning in Risk Analyses of Structural Engineering

Agreement Basics for Engineering

Advancements & Challenges in Fire Engineering Codes & Standards

View full program including keynote, receptions, exhibit hall and more at www.SEICON24.org

EDUCATION

FORENSIC

NATURAL DISASTERS

NON-BUILDING/SPECIAL STRUCTURES

1:45–2:45 PM

NON-STRUCTURAL

3:00 –4:00 PM Design of Safety-Related Structures

ASCE 7 Development, IBC Adoption and Changes in the 2022 Standard

NBIS & SNBI: Implementing the New National Bridge Inspection Standards & Specifications

Energy Transition Storage Challenges

Climate Change and Structural Loads

Business Issues Roundtable Discussion hosted by CASE ASCE at Work in Washington: The Biden Administration’s National Initiative to Advance Building Codes Considerations in Steel Modular Design and Fabrication

10:45–11:45 AM Sharing the Stories of Real Claims Innovations in Modular, Rapidly Erectable and Deployable Structures Case Studies of Complex Girder Erection & Temporary Work

TECHNOLOGY

4:45–5:45 PM

Engaging the Next Generation of Engineers

1:45–2:45 PM

RESEARCH

Performance-Based Design: Current State of the Art Blast: Challenging the Standard

The Customer Speaks Structural Engineering education from the eyes of current students & recent grads

3:00 –4:00 PM ASCE 41-23: A First Look Blast Hardening and Environmental Sustainability

Building a Better Structural Engineer - using improvisation to improve communication

NCSEA News Congratulations to the 2023 NCSEA Special Awards Honorees!

The NCSEA Special Awards are bestowed on individuals who exemplify outstanding service and commitment to the association and the structural engineering field.

This year's deserving honorees will be celebrated at the NCSEA Awards Celebration during the Structural Engineering Summit in Anaheim, CA. Congratulations to the outstanding recipients!

s n o i t a l u t a r g n Co

NCSEA Service Award The NCSEA Service Award is presented to an individual who has worked for the betterment of NCSEA, member organizations and the profession to a degree that is beyond the norm of volunteerism. It is given to someone who has made a clear contribution to the organization and, therefore, to the profession.

Robert Cornforth Award The Robert Cornforth Award is presented to an individual for exceptional dedication and exemplary service to a member organization and to the profession. The award is named for Robert Cornforth, a founding member of NCSEA treasurer on its first Board of Directors and a member of OSEA. Thomas A. Grogan Jr., P. E., S. E., has been honored with the NCSEA Service Award and the Robert Cornforth Award. Tom embodies what NCSEA needs – remarkable individuals who have given their all to both their SEA and NCSEA. His success illustrates what NCSEA is all about and is at the heart of what makes our profession great.

Tom has more than 40 years of experience within the structural engineering profession. His career began in 1980 in Washington, DC where he worked for Bechtel, GMR LTD. and Ellerbe Becket. In 2003, he moved to Jacksonville, Florida to work for Haskell until retiring in 2019 as Chief Structural Engineer. During his career Tom was registered as a Professional Engineer in 35 jurisdictions, including 5 as a Professional Structural Engineer.

He has been actively involved with FSEA, where he served on the Board of Directors from 2006 to 2014 and as President in 2010 and 2012. He has been Chair of the FSEA Licensure Committee since 2009, promoting SE licensure with the Florida legislature in several legislative sessions. His committee's efforts resulted in the passage of the Florida Structural Engineers Recognition Program in 2022.

Tom has been actively involved with NCSEA. He was a board member from 2012 to 2018, serving as President during 2016-2017. He has been a member of the NCSEA Licensure Committee since 2009 and has also worked closely with the Young Members Support Committee since its inception. Currently Tom serves as secretary/treasurer for the NCSEA Foundation.

Tom has been a member of ASCE (where he was awarded the title of Fellow), SEI, ACI, and AISC for most of his career and served on the NCEES SE Exam Committee from 2009 to 2013. He is very active with ACE mentoring in NE Florida where he recently completed a two-year term as Board Chair.

Susan M. Frey NCSEA Educator Award The Susan M. Frey NCSEA Educator Award is presented to an individual who has a genuine interest in and extraordinary talent for effective instruction of practicing structural engineers. The award was established to honor the memory of Sue Frey, one of NCSEA's finest educators.

Terry Malone, P. E., S. E., a licensed civil and structural engineer in the State of Washington and Senior Technical Director of Project Resources and Solutions Division of WoodWorks, Wood Products Council. Prior to joining WoodWorks, he was a principal in consulting structural engineering firms in Washington and Oregon and also conducted third-party structural plan reviews. He also served as a faculty member at St. Martin’s College (University) in Lacey, Washington. Terry has over 50 years of wood design experience and has taken an active role as a presenter, providing seminars at state and local ICC chapters and professional engineering organizations focusing on understanding lateral forceresisting systems. Terry is the author of The Analysis of Irregular Shaped Structures: Wood Diaphragms and Shear Walls, second edition, published by McGraw-Hill, ICC SKGA, and ICC.

follow @NCSEA on social media for the latest news & events!

58 STRUCTURE magazine

News from the National Council of Structural Engineers Associations James Delahay Award This award is presented at the recommendation of the NCSEA Code Advisory Committee to recognize outstanding individual contributions toward the development of building codes and standards. It is given in the spirit of its namesake, a person who made a long and lasting contribution to the code development process.

Thomas DiBlasi, P. E., is the president of DiBlasi Associates, P. C., in Monroe, CT. In his 38-year career as a structural engineer, he has been involved in a wide array of projects, including both new construction and renovation. He is a past president of NCSEA, where he served as chair of the Code Advisory Committee (CAC) from 2013-2022, and he has participated on CAC’s Wind Engineering Subcommittee since its inception. Mr. DiBlasi is past president of the Structural Engineers Association of Connecticut (SEAConn) and has served on its Code Advisory Committee for over 25 years. He has also served since 2005 as the structural engineering representative to the Codes & Standards Committee of the State of Connecticut, adopting and administering the Connecticut State Building Code in conjunction with the State Building Inspector and providing training to building officials.

Susan Ann “Susie” Jorgensen Presidential Leadership Award This award was created to honor the late NCSEA Board President and advocate for the profession, Susan Ann “Susie” Jorgensen. This award is presented to an individual who has demonstrated exceptional leadership potential through their activities within NCSEA and/or their SEA. The award is to be bestowed on candidates who embody Susan’s passion, vision, and legacy of leadership.

Marcus Freeman, P. E., has over 8 years of structural engineering experience and has worked on a wide variety of projects that include convention centers, airports, and high-rise residential. He received both his Bachelor’s and Master’s degrees in Structural Engineering from Virginia Tech and is currently a technical director at Taylor Devices in Seattle, WA. Marcus specializes in implementing fluid viscous dampers to improve the dynamic performance of structures. Marcus served as president of SEAW’s YMG (2016–2018) and currently serves as the chair for the Structural Engineering Engagement and Equity Committee (SE3).

NCSEA Webinars

Visit www.ncsea.com/education for the latest news on upcoming webinars and other virtual events.

November 16, 2023

Mastering Seismic Design: Answers to the Top 20 Frequently Asked Questions

November 28, 2023

Case Studies in Professional Ethics

December 14, 2023

An Introduction to the New ASCE Solar PV Structures Manual of Practice

Purchase an NCSEA webinar subscription and get access to all the educational content you’ll ever need! Subscribers receive access to a full year’s worth of live NCSEA education webinars (25+) and a recorded library of past webinars (170+) – all developed by leading experts; available whenever, wherever you need them!

Courses award 1.0 -1.5 hours of Diamond Review-approved continuing education after completing a quiz.

Recommendations for Performing Structural Engineering Quality Assurance Reviews

Recommendations for Performing Structural Engineering Quality Assurance Reviews

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SEI Update ASCE 7-22 Wind Loads Now Available Authors Stafford and Reinhold provide a comprehensive overview of the wind load provisions in Minimum Design Loads and Associated Criteria for Buildings and Other Structures, ASCE/SEI 7-22, focusing on providing direction while using the provisions that affect planning, designing, and constructing buildings for residential and commercial purposes.

Education

ASCE/SEI Substation Structure Design Guide-MOP 113 Program Join us for a discussion on the soon-to-be published ASCE/SEI Substation Structure Design Guide which will provide structural design guidance and function as a comprehensive resource for outdoor electrical substation structures and foundations. The live program on January 25 will include a dialogue with industry experts, a technical presentation and Q&A with the audience. Register today at www.asce.org/SEIEvents.

SEICon24 SEICon24 is right around the corner and we hope you are getting as excited as we are! Join us March 20-22 at NASCC: The Steel Conference for an exciting opportunity to learn, engage, and network with the vibrant structural engineering community. Be sure to check out the keynote address brought to you by David Odeh and featuring a trailer from ASCE’s latest movie, Cities of the Future.

Learn more at www.SEICon24.org

SEI Futures Fund Scholarships to SEICon24 Students & Young Professionals: Apply for an SEI Futures Fund scholarship to participate at SEICon24 at NASCC March 19-22 in San Antonio. Expand your horizons – learn the latest in structural engineering and connect with your SEI community! www.SEICon24.org/scholarships

Advancing the Profession

2024 SEI Futures Fund Initiatives Thank you for your generous giving! For 2024 the SEI Futures Fund Board has committed more than $200,000 for these strategic programs: • Student & Young Professional Scholarships to engage at in-person SEI conferences • Young Professional Travel Support on SEI Standards committees • SEI Carbon Impacts Workshop • Publish SEI Prestandard for Calculation Methodology for Structural Systems in Whole-Building Life Cycle Assessment • Advocacy of the ASCE 7-22 Supplement #2 for the new Flood Chapter Learn more and give to maximize the CSI 3 to 1 match at www.asce.org/SEIFuturesFund.

Follow SEI on Social Media: 60 STRUCTURE magazine

News of the Structural Engineering Institute of ASCE

Thank you to 2023 SEI Sustaining Organization Members Elite:

ALFRED BENESCH & COMPANY HARDESTY & HANOVER SCHNABEL FOUNDATION COMPANY SIMPSON STRONG-TIE WALTER P. MOORE

Join SEI as a Sustaining Organization Member to reach SEI members year-round, and show your support for SEI to advance and serve the structural engineering profession. www.asce.org/SEIMembership

Reinforced Aerated Autoclaved Concrete Roof Planks in the UK Widespread concerns over the safety of Reinforced Autoclaved Aerated Concrete (RAAC) roof planks in buildings in the UK have required safety measures to be put in place, including closures, to over 150 schools in England. Based on failures of structures in the UK employing such planks, the UK Standing Committee on Structural Safety (SCOSS) and Collaborative Reporting for Safer Structures – UK (CROSS-UK), both partners of SEI’s Collaborative Reporting for Safer Structures – US (CROSS-US), published an alert www.cross-safety.org/uk/safety-information/cross-safety-alert/failure-reinforced-autoclaved-aerated-concrete-raac-planks leading to extensive inspection of buildings in the UK that might contain such planks. The Washington Post covered this issue on September 5, 2023 at www.washingtonpost.com/world/2023/09/05/uk-raac-concrete-crisis-schools/

CROSS-US Call for Member Volunteers

Collaborative Reporting for Structural Safety – US (CROSS-US) cross-safety.org/us, an entity of SEI, helps professionals make structures safer by publishing safety information based on reports it receives and information in the public domain. CROSS-US’s Executive Committee (ExCom) seeks volunteers to serve as Members or Corresponding Members on one of its three subcommittees:

CROSS-US visibility development • Expand the audiences who should know about and participate in CROSS-US. • Develop methods for disseminating information: o Conferences o Webinars o Presentations o Articles and interviews o Continuing educational programs Report submission and development • In conjunction with the Visibility Development expand the audiences who should know about and participate in CROSS-US, as all these audiences are potential report submitters.

• Promote the use of the CROSS-US database as an educational tool and portal to educators and practitioners alike, demonstrating the value of report information and providing a pathway for additional report submissions.

• Make the process of submitting a report to CROSS-US as simple and straightforward as possible through a variety of social media postings, simple guidelines, and/or instructional videos.

Leadership and volunteer development • Expand leadership and volunteer resources to enable robust growth of CROSS-US over the next five years.

• Plan for long-term succession and engage younger members.

• Achieve diversity in leadership and volunteering.

• Involve individuals of broader backgrounds.

• Engage key stakeholders and potential sponsors.

It is anticipated that each subcommittee will meet on-line approximately monthly, and that members and corresponding members will do a modest amount of task and preparatory work between meetings. Selection criteria include (1) ability and interest to commit to CROSS-US’s and the subcommittee’s mission, (2) ability to collaborate effectively in groups, and (3) strong oral and written communication skills.

To apply, state your interest in which particular committee, indicate your response to selection criteria above, and provide a brief bio including current member #, educational background, professional experience, other related activity with SEI/ASCE or other organizations, and direct to Suzanne Fisher at sfisher@asce.org for CROSS-US ExCom consideration.

Errata

SEI Standards Supplements and Errata including ASCE 7.

See www.asce.org/SEI. To submit errata, contact sei@asce.org.

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CASE in Point Tools To Help Your Business Grow...

CASE has committees that work together to produce specific resources available to members, from contract documents to whitepapers, to help your business succeed.

If you are a member of CASE, all CASE publications are free to you. NCSEA and SEI members receive a discount on publications. Use discount code - NCSEASEI2022 when you check out.

Check out some of the brand new CASE Publications developed by the Guidelines CommitteeÉ CASE 976-C: A Review and Commentary on the American Institute of Steel Construction 2022 Code of Standard Practice for Steel Buildings and Bridges The importance of the AISC Code of Standard Practice (AISC 303-22), referred to herein as the Code or COSP) to the construction community is manifested in its almost 100 years of use and development.

This Code establishes the trade practices for the steel industry. Generally, this involves the acceptable practices and responsibilities of the Fabricator and Erector and the responsibilities of others such as the Owner’s Designated Representative for Design (ODRD) – (usually the Structural Engineer of Record), the Owner and the Owner’s Designated Representative for Construction (ODRC) – (usually the General Contractor or Construction Manager or similar authority at the jobsite) as they relate to the work of the Fabricator and Erector. The 2022 COSP addresses many recent changes in the practice of designing, purchasing, fabricating, and erecting structural steel and is therefore a continuation of the trend of past improvements and developments of this standard.

CASE White Paper Beyond the Code: Shrinkage Cracking

CASE recognizes that the International Building Code or other governing codes do not address all aspects of structural engineering and design. Often, the most common issues where the owners, or the contractor or the design team are not aligned deal with what is not clearly addressed by the various codes or design guidelines. This is the second in a series of “Beyond the Code” white papers that will attempt to collate design considerations that need to be discussed with the owners at the beginning of a project to establish a clear Basis-of-Design for the project. By proactively bringing up the design consideration in front of the owners, the Structural Engineer can set up realistic expectations and discuss the cost impact of alternative designs.

This white paper in the “Beyond the Code” series discusses shrinkage cracking in concrete with an explanation of why it occurs, common locations they occur, and strategies to mitigate them becoming a risk in your project.

You can purchase these and other Risk Management Tools at www.acec.org/bookstore.

You can also browse all of the CASE publications at www.acec.org/coalitions/coalition-publications/ Is there something missing for your business practice? CASE is committed to publishing the right tools for you.

Have an idea? We’d love to hear from you!

Follow ACEC Coalitions on LinkedIn: www.linkedin.com/in/acec-coalitions 62 STRUCTURE magazine

News of the Coalition of American Structural Engineers Upcoming Events Joint Town Hall Event with CASE, NCSEA, and SEI February 21, 2024 2:00–3:30 pm ET Online Leadership from CASE, NCSEA, and SEI will host a virtual joint town hall event to discuss how the three organizations are progressing to fulfill the Vision for the Future of Structural Engineering (adopted April 2019), highlighting initiatives to advance the profession and enhance member engagement.

The town hall is an opportunity to catch up on things you might have missed and gain insight into what the three organizations are doing moving forward.

This complimentary event is open to all CASE, NCSEA and SEI members. https://program.acec.org/ joint-town-hall-event-case-ncsea-and-sei

Generation ZE: Transitioning to Zero Emissions November 14, 2023 1:30–2:30 pm ET Online Hosted by ACEC’s Coalition of American Mechanical and Electrical Engineers, join in the conversation about Zero Emission trends and its far-reaching impact on the mobility sector. https://www.acec.org/event/ generation-ze-transitioning-to-zero-emissions/

Now more than ever we need to support the upcoming generation of the workforce.

Give to the CASE Scholarship today!

N OVE M B ER 2023

63

INSIGHTS HSS Availability

Checking the availability of HSS sections is easy with the Steel Tube Institute’s Capability Tool.

By Beth Suminski, P. E., S. E., and Cathleen Jacinto, P. E., S. E.

I

f you have ever opened the steel manual to find section properties for a steel HSS (hollow structural section) member, you may have been surprised by how many sizes of HSS are listed. There are hundreds of options for round, rectangular, and square tubes. Some of these sections are readily available, others are produced on demand, and some may not be available in the current market. Realizing this, the Steel Tube Institute (STI), an industry organization supported by domestic steel tube producers, has created a free Capability Tool as a resource to search which HSS sections are currently domestically produced, by whom, and whether the sections are regularly produced or produced on demand.

At steeltubeinstitute.org/capability-tool, there are search fields for shape, dimension range, and grade. Three ASTM (American Society for Testing

and Materials) HSS material grades are included: A500, A1085, and A1065; and one for mechanical tubing, A513. Once these fields are filled, a list of producers who provide that shape, either regularly or on demand, is generated.

ASTM A500 is the current market leader for HSS material. It is widely available in many sizes, shapes, and wall thicknesses in North American service centers. For this reason, the availability tables in this article are based on A500 material. Domestically produced A500 HSS members are dual certified as grades B/C. The B/C dual certification means that the material meets the requirements for both grades. Therefore, when utilizing A500 HSS, STI recommends specifying and designing with Grade C, which has strengths of Fy = 50 ksi and Fu = 62 ksi for squares, rectangles, and round

Table 1 Readily Available ASTM A500 Square HSS Sections [2 or More STI HSS Producers]

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shapes, per ASTM A500-21.

Three other material grades that are commonly specified for steel tubes have some overlap with A500 but also provide unique properties of their own. A1085 is a newer HSS material grade with tighter wallthickness tolerances, additional Charpy V-notch requirements, and strengths of Fy = 50 ksi and Fu = 65 ksi. A1085 HSS is only produced on demand and is not readily available in the current market. A1065, another HSS material grade, provides similar strength characteristics to A500 and A1085 but is manufactured differently, allowing tapered sections, multi-sided sections (i.e., lighting poles), and larger square and rectangular sizes to be formed. Round sections are not produced using A1065. A513 is a mechanical tubing specification that is not structurally graded. It is intended for applications requiring strict dimensional tolerances, but where the member’s strength is not critical. If using A513 for structural applications (i.e., handrails), its strength MUST be verified through reports or testing.

You may have noticed the omission of a material specification, ASTM A53, often utilized for round sections. A53 is a mechanical pipe specification, not a structural specification, and there are good reasons for that. A53 focuses on producing a pipe that is rigorously tested to withstand pressurized conditions and that is coated in an exterior sealant, among other characteristics. The costs of these additional piping requirements are passed along even in non-piping structural usage. A53’s yield strength is also less than the structural grades of A500 and A1085.

For these reasons, round structural tubular sections should be specified as A500 or A1085, not A53. If you find yourself asked to replace A500 round HSS with A53 pipe or vice versa, this article with a list of round HSS sizes that match common pipe sizes is helpful: steeltubeinstitute. org/resources/selecting-rightround-hss/. But remember that if A53 is ultimately used for structural applications, its strength MUST be verified through reports or testing.

Table 2 Readily Available ASTM A500 Rectangular HSS Sections [2 or More STI HSS Producers] N OVE M B ER 2023

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Table 3 Readily Available ASTM A500 Round (<=4” Diameter) HSS Sections [2 or More STI HSS Producers]

Table 4 Readily Available ASTM A500 Round (>=4.5” Diameter) HSS Sections [2 or More STI HSS Producers]

The next time you are faced with hundreds of HSS choices, we encourage you to utilize the Steel Tube Institute’s free Capability Tool for the most up-to-date HSS section availability. A snapshot of production at the time this article was written is given in the included tables of commonly available square, rectangular, and round HSS A500 Grade B/C sections from multiple domestic HSS producers. A dot in Tables 1 through 4 indicates that as of August 2023, the corresponding HSS section was readily available from 2 or more STI HSS producers based on regular HSS production rolling schedules. If a section you wish to specify is not listed in the tables, it does not mean it is unavailable. Individual suppliers regularly roll sizes that are not indicated here. Ultimately, the regularly-updated Capability Tool is the best source of available sizes from each STI HSS producer, providing a more comprehensive selection beyond what is presented here.■ Beth Suminski and Cathleen Jacinto serve as HSS Technical Consultants to the Steel Tube Institute and are practicing Structural Engineers with FORSE Consulting. (hssinfo@steeltubeinstitute.org).

Don’t be overwhelmed by the hundreds of HSS choices.

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